thin-film thermo-mechanical sensors embedded in metallic structures

Transcription

THIN-FILM THERMO-MECHANICAL SENSORS
EMBEDDED IN METALLIC STRUCTURES
A DISSERTATION
SUBMITTED TO THE DEPARTMENT OF
MATERIALS SCIENCE AND ENGINEERING
AND THE COMMITTEE ON GRADUATE STUDIES
OF STANFORD UNIVERSITY
IN PARTIAL FULFILLMENT OF THE REQUIREMENTS
FOR THE DEGREE OF
DOCTOR OF PHILOSOPHY
Anastasios Golnas
December 1999
© Copyright by Anastasios Golnas 2000
All Rights Reserved
ii
I certify that I have read this dissertation and that in my opinion
it is fully adequate, in scope and quality, as a dissertation for the
degree of Doctor of Philosophy.
Friedrich B. Prinz (Principal Adviser)
William D. Nix
David M. Barnett
Approved for the University Committee on Graduate Studies:
_______________________________________________
Dean of Graduate Studies
iii
Abstract
The ability to monitor in real time the thermo-mechanical responses of tools,
equipment, and structural components has been very appealing to the aerospace,
automotive, drilling, and manufacturing industries. Such responses can be
measured with the appropriate sensors: thermocouples for temperature and
strain gages for deformation. So far, the challenge has been to instrument the
tools, equipment, or structural components with a number of sensors in an
economical way and also protect the sensors from the environment which the
tools, etc. are exposed to.
In this work, a sequence of manufacturing processes that can be used to build
thin-film temperature and strain sensors on internal surfaces of metallic
structures is proposed and demonstrated. The use of thin-film techniques allows
the parallel fabrication of sensor arrays, whereas a layered manufacturing
scheme, e.g.: Shape Deposition Manufacturing, permits the creation of sensors on
the internal surfaces of metallic parts and their subsequent embedding as the
parts are completed.
Specifically, thin-film sensors are deposited on an insulating aluminum oxide
film, which is grown on a polished stainless steel substrate. The oxide is
deposited by sputtering an aluminum target in a reactive atmosphere. The
sensors are sputter-deposited from alloy targets, shaped via micromachining and
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partially covered with a passivation layer of aluminum oxide. The thin-film
structure is then partially covered by two protective layers of copper and nickel.
Both layers are electroplated and their purpose is to protect the thin films during
the deposition of the embedding layers. Embedding is accomplished by using a
high-power infrared laser to melt a powder bed on top of the protective layers.
Invar, a Fe-Ni alloy with very small coefficient of thermal expansion, is used as
embedding material in order to minimize the effect of residual thermal stresses
on the structure.
Among the issues that emerged during the definition of the fabrication sequence
were: the long-term stability of reactive deposition, the presence of pinholes in
the dielectric layers, the adhesion of the thin films to their substrates, the
reactions between the thin films and the electroplating bath, the adhesion of the
electroplated layers to the substrate, the optimal combination of materials and
thickness of the protective layers, the bonding between the embedding and the
protective layers, and the heat input and residual stresses resulting from the
high-temperature embedding process. Last but not least, the calibration of the
deposited thin-film sensors and their splicing to measuring equipment had to be
addressed.
Finally, a finite element model was constructed in order to simulate the hightemperature embedding process. The heat transfer analysis performed on the
model provides the temperature profiles of all nodes and can be used as a tool for
the optimization of the protective layer thickness. Its results can also be used for
a stress analysis of the multilayered structure.
In conclusion, an integration of technologies is offered to allow the
instrumentation of metallic parts with embedded thermo-mechanical sensors
during the manufacturing sequence, in an environment without stringent
particle control, while making use of commonly available materials.
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Acknowledgments
As the formal completion of my doctoral studies in the Department of Materials
Science and Engineering is under way, I am very pleased to have the honor to
express my gratitude from the pages of my dissertation to all those who were
involved in this journey in research and applied knowledge.
Initially, I would like to thank my thesis adviser, Professor Fritz Prinz, who gave
me the opportunity to join his group in the spring of 1995. I am still grateful for
having been offered a chance to gain direct engineering knowledge and apply
my physical intuition by joining a research group where the focus is the
fabrication of novel structures and devices. I also feel gratitude for his
continuous support and encouragement that helped me overcome the
disappointments that are so abundant in any experimental and engineering
problem. Finally, I must thank him for our numerous discussions on a variety of
scientific, engineering, cultural, and political subjects which augmented my
vision and understanding.
I am also indebted to Professor Robert Merz who was instrumental in the
conception of the embedded sensors research project and the realization of the
mesoscopic structures laboratory. Robert provided the early framework of the
project, recruited me as the student in charge, and served as an excellent guide in
the first stages of the research.
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It is my honor to acknowledge the support of Professors David M. Barnett and
William D. Nix who served in my Reading and Defense Committees, as well as
Professor Martin Fischer who chaired my Defense Committee. I owe special
thanks to Professor Michael Kelly, also a member of my Defense Committee, who
was an excellent resource in my quest for the cause of pinholes in the dielectric
thin films, both through his class and his personal consulting.
I would also like to thank the members of my Qualifying Examination
Committees: Professors W.D. Nix, R. Sinclair, R.H. Dauskardt, and T. Huffnagel.
Furthermore I am grateful to the Department, its faculty, students, and
administration for providing a superb environment for learning, and for offering
their help in every occasion I needed it. It has been an unmatched experience.
I owe many thanks to my colleagues in the Rapid Prototyping Laboratory (RPL)
for their help with my research and preparation for the Oral Examination, as well
as for their support all these years. I am especially indebted to Tom Hasler, Dr.
Jürgen Stampfl, and Rudi Leitgeb, who helped with various aspects of the
deposition chamber operation. Special thanks are due to my officemates, Dr.
Alexander Cooper, Dr. Alex Nickel, and Dr. John Kietzman, with whom I
engaged in vastly entertaining discussions. I also want to thank Sylvia Walters
and Lynn Hoschek for their administrative help throughout my service in the
RPL.
All research on the embedded sensors project was funded by the Office of Naval
Research through the contract N00014-96-1-0354-P00003, and it is my satisfaction
to acknowledge their continuing support for a second three-year period.
Concluding my expression of gratitude to the scientific community at Stanford, I
would like to make a very special reference to my colleagues Xiaochun Li of the
RPL, and Kevin Ohashi and Dr. Surya Iyer of the R.H. Dauskardt group. Their
contribution to the work presented in the next pages cannot be overstated. Surya
(now with Applied Materials, Inc.) did all the setup for the first experimental
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strain measurements of the thin film strain gages, in March of ’98. Kevin
contributed many hours helping me to repeat those experiments, twice, in the
summer of ’99. Without their input it would have been very difficult to test the
sensors. Xiaochun’s contributions were also of central importance to the project.
Apart from his help with the sample preparation, he operated and troubleshot
the laser deposition station for the embedding experiments, and offered excellent
insight in many stages of the project and especially in the high-temperature
aspects. I am also indebted to Professor Reinhold H. Dauskardt for his kind
permission to use the testing equipment in his laboratory, Ali Farvid from SLAC
for his help in establishing an electroplating facility, and Bob Jones from CMR for
his help with the EPMA measurements.
It would be an omission not to mention the people who inspired, encouraged,
helped, and allowed me to come to one of the premier research institutions of
this country. Namely, I want to thank my senior thesis adviser Professor Stergios
Logothetidis, and Professors D. Gounaris, G. Theodorou, N. Economou, E.
Paloura, K. Paraskevopoulos, and E. Hatzikraniotis, all from the Physics
Department of the Aristotle University in Thessaloniki, Greece. I also want to
thank the admissions committee of the Department of Materials Science and
Engineering, Professor David M. Barnett who served as my academic advisor
during my first year at Stanford, and the School of Engineering for the
Fellowship that supported me during the first academic year.
Last, but not least, I would like to express my gratitude to all those who made me
feel at home with their love, friendship, and care, and especially to Athina
Vassilakis, Dimitris Pantelidis, Constantinos Papadias, Stelios Diamantidis,
Phaedon Kyriakidis, Raj Vaidyanathan, Joe Tringe, and Fr. Peter Salmas. Of
course, I will always be indebted to my old friends at home and abroad. Finally, I
would like to express my deep gratitude and love to my family and acknowledge
their unconditional trust and multifaceted support through all these years.
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Table of Contents
Abstract
iv
Acknowledgments
vi
Table of Contents
ix
List of Tables
xii
List of Figures
xiv
1 Introduction
1
1.1 Incentive . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1
1.2 Thin-film Sensors . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3
1.2.1 Applications . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3
1.2.2 Manufacturing Issues . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5
1.3 Basics of Sensor Physics . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7
1.3.1 Strain Sensors. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7
1.3.2 Thermocouple Sensors . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9
1.4 Deposition Processes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11
1.4.1 Sputtering . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11
1.4.2 Electroplating. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21
1.4.3 Laser Assisted Deposition . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 23
1.5 Dissertation Outline. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 24
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2 Proposed Solution
25
2.1 Cleanliness and Surface Quality . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 25
2.2 Insulation. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 26
2.3 Sensors . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 26
2.4 Thermal Protection . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 27
2.5 Embedding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 27
3 Fabrication
28
3.1 Substrate preparation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 28
3.1.1 Cutting . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 28
3.1.2 Grinding and polishing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 29
3.1.3 Cleaning . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 30
3.1.4 Sputter-etching . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 31
3.1.5 Discussion of cleaning processes . . . . . . . . . . . . . . . . . . . . . . . . . . 33
3.2 Bottom insulation layer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 34
3.2.1 Determination of transition oxygen flow . . . . . . . . . . . . . . . . . . . 36
3.2.2 Deposition sequence . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 49
3.3 Sensor layer. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 53
3.3.1 Photolithography . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 53
3.3.2 Deposition . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 55
3.3.3 Lift-off . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 56
3.4 Top Oxide . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 57
3.5 Protective Layers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 58
3.5.1 Copper layer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 60
3.5.2 Nickel layer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 63
3.6 Embedding Layers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 65
3.6.1 Thermal stress minimization . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 66
3.6.2 Fusion of the powder layer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 67
3.7 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 71
4 Characterization
73
4.1 Composition . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 73
x
4.1.1 Dielectric layers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 73
4.1.2 Sensor layers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 77
4.2 Microstructure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 79
4.2.1 Dielectric layers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 79
4.2.2 Sensor films . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 81
4.2.3 Protective layers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 81
4.3 Electrical measurements . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 83
4.3.1 Dielectric strength of dielectric layers . . . . . . . . . . . . . . . . . . . . . . 83
4.3.2 Resistivity of sensor materials. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 85
4.4 Strain-gage characterization . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 87
4.4.1 Experimental setup . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 87
4.4.2 Experimental objective and results. . . . . . . . . . . . . . . . . . . . . . . . . 89
4.5 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 98
5 Modeling
100
5.1 Model Description . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 100
5.1.1 Geometry . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 100
5.1.2 Material Properties . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 104
5.1.3 Simulation of a moving heat source . . . . . . . . . . . . . . . . . . . . . . . 106
5.1.4 Boundary and initial conditions . . . . . . . . . . . . . . . . . . . . . . . . . . 107
5.2 Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 110
5.3 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 114
6 Conclusions
116
A The thin-film deposition system
120
B Positive masks of strain sensors
124
C Sample LabView file
125
D ABAQUS heat transfer input file
128
Bibliography
142
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List of Tables
Table 3.1:
Nominal composition of 304L stainless steel. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 29
Table 3.2:
Sputter-etching process parameters. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 32
Table 3.3:
Surface composition of 304L substrates cleaned with different methods. . . . . . . . . . 33
Table 3.4:
Atomic content of O, Fe, and Cr in 304L surfaces after sputter-etching. . . . . . . . . . . 34
Table 3.5:
Material properties for insulating thin films. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 35
Table 3.6:
Linear fit parameters for high oxygen flows (Figure 3.1 and Figure 3.2). . . . . . . . . . 39
Table 3.7:
Characteristics of the voltage-flow hysteresis curves in Figure 3.1. . . . . . . . . . . . . . . 40
Table 3.8:
Process parameters for reactive deposition of aluminum oxide. . . . . . . . . . . . . . . . . 50
Table 3.9:
Deposition parameters for the sensor alloys and the adhesion interlayer. . . . . . . . . 56
Table 3.10: Deposition parameters for the “seed” copper film and the adhesion interlayer. . . . 61
Table 3.11: Chemical composition of copper electroplating solution. . . . . . . . . . . . . . . . . . . . . . . 63
Table 3.12: Electrodeposition parameters for the copper protective layer. . . . . . . . . . . . . . . . . . . 63
Table 3.13: Chemical composition of nickel “strike” electroplating solution. . . . . . . . . . . . . . . . 64
Table 3.14: Electrodeposition parameters for the “seed” nickel layer. . . . . . . . . . . . . . . . . . . . . . . 64
Table 3.15: Chemical composition of nickel sulfamate electroplating solution. . . . . . . . . . . . . . . 64
Page xii
Table 3.16: Electrodeposition parameters for the nickel protective layer. . . . . . . . . . . . . . . . . . . . 65
Table 4.1:
Nominal composition of constantan, measured values for deposit. . . . . . . . . . . . . . 78
Table 4.2:
Nominal composition of chromel, measured values for deposit. . . . . . . . . . . . . . . . . 78
Table 4.3:
Nominal composition of alumel, measured values for deposit. . . . . . . . . . . . . . . . . . 78
Table 4.4:
Resistivity and thickness of alumel, chromel, and constantan films. . . . . . . . . . . . . . 86
Table 4.5:
Dimensions and moment of inertia for beam substrate. . . . . . . . . . . . . . . . . . . . . . . . 87
Table 5.1:
Nodal density along the three axes. Initial model. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 103
Table 5.2:
Nodal density along the three axes. Final model. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 103
Table 5.3:
Density and thermal properties of materials in the FE model. . . . . . . . . . . . . . . . . . . 105
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List of Figures
Figure 1.1: Generic hysteresis curve in a reactive sputtering system.. . . . . . . . . . . . . . . . . . . . . 17
Figure 1.2: Asymmetric bipolar dc voltage as applied to a magnetron target. . . . . . . . . . . . . . 20
Figure 3.1: Discharge voltage and total pressure vs. oxygen flow for two Al targets. . . . . . . . 37
Figure 3.2: Variation of total pressure with oxygen flow. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 38
Figure 3.3: Transition voltage and flow for two Al targets (Figure 3.1 & Table 3.7).. . . . . . . . . 40
Figure 3.4: a) Transition flow and power values for depositions from one Al target.
b) Transition flow vs. corresponding power level (from graph a). . . . . . . . . . . . . . 41
Figure 3.5: Actual (markers) and predicted (line) transition flows for 4 depositions. . . . . . . . 43
Figure 3.6: a) Calculation of partial pressure–oxygen flow relation.
b) Total pressure vs. increasing oxygen flow with and without discharge. . . . . . . 46
Figure 3.7: Oxygen consumption rate vs. discharge power.. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 48
Figure 3.8: Shadow mask for shaping sputter-deposited Al2O3 layers. . . . . . . . . . . . . . . . . . . 49
Figure 3.9: Discharge voltage and oxygen flow variations during reactive deposition. . . . . . 52
Figure 3.10: Schematic of the photolithography steps preceding a lift-off process. . . . . . . . . . . 54
Figure 3.11: Jagged edge of deposited constantan film after lift-off.. . . . . . . . . . . . . . . . . . . . . . . 57
xiv
Figure 3.12: Cross-section showing an Al2O3 film damaged during laser deposition. . . . . . . . 59
Figure 3.13: a) Clustered voids at the invar/copper interface.
b) Distinct phases across the partially remelted invar/copper interface.. . . . . . . . 60
Figure 3.14: a) Schematic of substrate with sensor enclosed between two dielectric films.
b) Schematic showing the copper layer footprint.
c) Schematic showing the nickel layer footprint. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 60
Figure 3.15: Laser path during progressive attempts to improve embedding. . . . . . . . . . . . . . . 68
Figure 3.16: Invar-nickel interface from a cross-section of an embedded structure. . . . . . . . . . 69
Figure 3.17: Final laser path configuration.. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 70
Figure 4.1: Oxygen-to-aluminum ratio vs. sputter-etching time. . . . . . . . . . . . . . . . . . . . . . . . . 75
Figure 4.2: Aluminum content of deposited Al2O3 film vs. sputter-etching time.. . . . . . . . . . 76
Figure 4.3: SEM image of amorphous Al2O3 film on a stainless steel substrate. . . . . . . . . . . . 80
Figure 4.4: XRD spectra of Al2O3 on 304L, and uncoated 304L substrate. . . . . . . . . . . . . . . . . 81
Figure 4.5: Grain sizes in sputtered and electroplated copper layers. . . . . . . . . . . . . . . . . . . . . 82
Figure 4.6: Copper grains in electroplated layer. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 82
Figure 4.7: Optical micrograph of damage caused by self-healing dielectric breakdown. . . . 84
Figure 4.8: SEM image of Al2O3 film with pinholes. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 85
Figure 4.9: Dimension details of the 4-point bend test setup. . . . . . . . . . . . . . . . . . . . . . . . . . . . 89
Figure 4.10: Outer fiber stress and of deposited strain gage output from 4-point bend test. . . 90
Figure 4.11: Displacement and calculated outer fiber stress as functions of time. . . . . . . . . . . . 91
Figure 4.12: a) Experimental gage output and nominal strain vs. calculated strain and stress.
b) Nominal strain vs. experimental gage output.. . . . . . . . . . . . . . . . . . . . . . . . . . . . 92
Figure 4.13: a) Calculated outer fiber stress and output of deposited strain gage.
xv
b) Calculated outer fiber stress and strain measured by calibrated strain gage. . . 94
Figure 4.14: a) Outer fiber stress vs. experimental gage output.
b) Outer fiber stress vs. nominal strain. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 95
Figure 4.15: Maximum stress and experimental gage response prior to plastic deformation. . 96
Figure 4.16: a) Stress-strain curves for the experimental sensor.
b) Stress-strain curves for the commercial sensor. . . . . . . . . . . . . . . . . . . . . . . . . . . . 96
Figure 4.17: Nominal strain vs. deposited sensor output. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 97
Figure 5.1: Materials and dimensions of the 4 strata in the geometric model. . . . . . . . . . . . . 102
Figure 5.2: Quantization of the powder bed in elements, top view. . . . . . . . . . . . . . . . . . . . . . 107
Figure 5.3: Heat flux for element sets 1 and 2 in Figure 5.2 and total heat flux. . . . . . . . . . . . 108
Figure 5.4: Experimental and calculated temperature at a point on the Ni-invar interface. . 111
Figure 5.5: Calculated maximum temperatures at the Ni-invar interface.. . . . . . . . . . . . . . . . 112
Figure 5.6: Calculated temperature at the steel-copper interface. . . . . . . . . . . . . . . . . . . . . . . . 113
Figure B.1: The numbers denote the smallest linewidth in µm.. . . . . . . . . . . . . . . . . . . . . . . . . 124
xvi
1
Introduction
The goal of this chapter is to provide the necessary background for the
development of the dissertation in the sections that follow. Particularly, an
incentive for the embedding of thermo-mechanical sensors is presented, followed
by a review of the recent progress on thin-film sensors. In addition, a brief
reference to the physics of thermo-mechanical sensing is made as well as a
description of the processing techniques that are useful in making and
embedding the sensors.
1.1 Incentive
The idea for embedded sensors came forth by the combination of a commercial
need and the existence of an enabling manufacturing concept. It was understood
that in various industrial practices it is useful to know in real time the basic
operating conditions (i.e.: temperature and deformation or strain) of certain
mechanical tools and components. Furthermore, it could be desirable to know
these conditions over extended areas of those parts. Examples of cases in which
such a concept is useful include the manufacturing industry (molds, dies, drilling
bits, etc.), the aerospace industry (components of jet engines), the oil industry
(drilling equipment), the power industry (vessels and pipes), the automotive
1
CHAPTER 1: INTRODUCTION
industry (components of motors), as well as the construction industry (structural
components in buildings).
Temperature and strain information can only be obtained by placing sensors in
those tools, and information from extended areas can only be obtained from
arrays of such sensors. Such a solution would call for the placement of the
sensors near the points of interest, and therefore the issues of assembly and
protection would have to be successfully addressed. The assembly of a large
number of sensors is cumbersome, time-consuming, and costly, and this
endeavor might become very difficult for tools operating in harsh environments.
However, the layered manufacturing of components and tools could allow for
the placement of sensors close to the points of interest and for their subsequent
enclosure. Besides, the use of Large Scale Integration (LSI) thin-film processes for
the fabrication of the sensors, similar to the ones used in the Integrated Circuit
industry, could allow for the simultaneous fabrication of sensor arrays. Thus, the
combination of these methods would offer proximity to the desired places,
protection from the environment during operation, and the large scale
integration of units (i.e. sensors) without the need for extensive postmanufacturing assembly [merz97].
The obvious challenges for this idea result from the fact that most tools and
components in the manufacturing, automotive, power, and oil industries, are
metallic. Therefore, the sensors (which are essentially simple passive circuits)
should be adequately insulated from the surrounding conductive matrix.
Furthermore, any layered fabrication technique that is designed to produce
functional metallic parts has to add the layers in a high-temperature state. This
will achieve a high-quality interlayer bonding via either interdiffusion or partial
remelting of the substrate surface. In this case, the sensors, and their insulation,
will have to be protected during the high-temperature deposition steps.
2
Last, but not least, comes the problem of the successful integration of two
manufacturing techniques that are fundamentally different in terms of scale and
production environments. In most cases, in IC fabrication lines, only silicon and a
small number of carefully selected elements are processed in order to produce
exceedingly small structures. On the other hand, in large-scale manufacturing
plants, where traditional engineering materials are heavily or exclusively used,
the conditions are unsuitable for thin-film processing, which needs a very
controlled environment.
1.2 Thin-film Sensors
1.2.1 Applications
Surface thermocouples have been produced by thin-film techniques since the
1930s by vacuum evaporation [burg30], or sputtering [harr34]. In the more recent
past, the aerospace industry has intensively explored the use of thin-film
thermocouples (TFTCs) as well as strain gages (TFSGs) and flux meters for
measuring the conditions on jet-engine turbine blades.
Such work was done at Pratt & Whitney Aircraft in the early 1970s, where TFTCs
suitable for high temperatures (Type S: Pt/Pt0.9Rh0.1) were sputter-deposited on
iron superalloy (FeCrAlY) substrates [gran85]. The substrates were oxidized in
air to form an insulating layer of aluminum oxide prior to the deposition of the
sensor films.
During the same period NASA began contracting research on thin-film sensors,
and researchers in United Technologies Corp. built TFTCs and TFSGs on typical
turbine blade materials (MAR-M-200 with Hf, Hastelloy X, and B1900 with Hf).
The substrates were first coated with a nickel superalloy layer (NiCoCrAlY) that
was thermally treated in a controlled atmosphere in order to grow a surface
3
aluminum oxide film. In later experiments that were conducted mostly at the
NASA Lewis Research Center, an additional alumina film was deposited by
evaporation or rf-sputtering from a compound target [mart94, lei97]. In yet
another implementation of TFTC fabrication a “seedcoat” of aluminum oxide
was sputtered on the superalloy layer before the substrate was oxidized in a
vacuum furnace with 2% O2 using nitrogen and hydrogen as forming gases
[nasa86]. The materials used to create the sensors were Pt and 90Pt-10Rh for the
thermocouples (Type S) and NiCr or PdCr for the strain gages. Type S materials
are traditionally used in high-temperature environments (up to ~1100°C), while
PdCr (13%wt. Cr) was selected as the best candidate for high-temperature strain
measurements [huls87]. In order to further protect the sensors against corrosion,
an aluminum oxide overcoat was deposited either by rf-sputtering or by
evaporation onto the metal films [mart94].
Researchers at the French National Institute for Aerospace Research and Studies
(ONERA) have been building thin-film thermocouples [gode87] and flux meters
[gode90] since the 1980s, as well as strain gages [kays93]. The fabrication
sequence has essentially the same components with the one developed by NASA
and its contractors: Type-S TFTCs, NiCr or PdCr TFSGs, or platinel-type
fluxmeters were deposited on ~8µm of Al2O3 which was also grown on
substrates made of commercial nickel-base superalloys with a NiCoCrAlY
coating (~30µm). All films and coatings were deposited by rf-sputtering of alloy
or compound targets. Finally, the sensors were covered with a thin (0.5µm)
protective layer of Al2O3 or SiO2.
A strong motivation behind efforts to detect the temperature pattern on the
surface of turbine blades comes from the need to design blades that withstand
ever higher combustion temperatures. As the operating temperatures are
elevated, first stage superalloy blades have to be cooled in order to survive. Film
cooling through discreet holes or slots on the blade surface has been selected as a
most effective method since the 1950s. However, it is very difficult to
4
theoretically predict the effectiveness of film cooling, therefore it has to be
determined experimentally [simo93, thak99]. Most research groups have
employed stationary flat or mildly curved plates to study the effect and data
were obtained by wire thermocouples [sinh91, jump91]. It seems that arrays of
TFTCs deposited on the surface of the turbine blade provide a more unobtrusive
and efficient way to gather temperature data at a high spatial resolution [fess99].
Thin-film sensors have also been used to monitor the performance of pressure
transducers. The latter are widely used in a large variety of applications ranging
from the chemical processing industry to the automobile engines. Researchers
from the Indian Liquid Propulsion Systems Centre and the Indian Institute of
Science deposited a manganese sensing film on the steel diaphragm of a pressure
transducer [naya90]. The diaphragm had previously been coated with an
insulation multilayer consisting of alternating Al2O3 and SiO2 films. The sensor
was finally covered with a 0.2µm SiO2 film. All layers were deposited by
evaporation and the insulating films were deposited in a reactive environment.
1.2.2 Manufacturing Issues
Substrate preparation
In almost all the examples referenced in “Applications” on page 3, substrate
preparation was the first key issue in the fabrication sequence. Its importance
became apparent through its relation to the adhesion and quality of the
insulating layer [naya90, garc93, stor96, nisk98]. The preparation techniques
involved polishing of the metallic substrate to a mirror surface finish and
subsequent cleaning with degreasing agents. Most of this processing was done in
clean-room environments (class 1000-100). Some groups employed sputteretching of the substrates prior to the insulator deposition as a final cleaning step.
A more thorough investigation of the preparation process was carried out by a
Spanish research team [garc93] on 17.4 PH stainless steel substrates. They
5
concluded that chloride particles attached to the surface were the main culprits
for defects in the overdeposited dielectric layer as well as for corrosion-induced
holes on the steel surface. They also observed that alkaline cleaners (dissolved
NaOH) would remove such particles and that there is a definitive increase of the
breakdown voltage of the dielectric layer for smoother substrate surfaces and
thicker dielectric films. Their suggested cleaning method takes place in a class
1000-100 clean room and consists of 12 distinct steps.
Deposition of insulating films
The deposition of defect-free, adherent insulating films has been the single
greatest challenge for creating thin-film sensor devices on metallic surfaces. Most
groups employed radio frequency (13.56MHz) sputtering of a compound target
as the deposition method of choice. The insulating film was aluminum oxide for
all the work done by the aerospace industry teams and a combination of
aluminum oxide and silicon dioxide for the team from India.
Radio frequency (rf) sputtering is a very slow deposition process and the
reported growth rates were 0.03-0.04nm/s. Since a film thickness of 6-8µm was
deemed necessary for adequate insulation, the total deposition time was on the
order of 50 hours [gode87, lei97].
In many cases the substrate was heated (200-900°C) during the aluminum oxide
deposition, in order to aid the surface mobility of the adsorbed species, promote
the formation of the metastable γ-phase (which is more corrosion resistant than
the amorphous), and cause the development of residual compressive stresses in
the ceramic film upon cooling of the system to room temperature. The last effect
would permit the oxide film to remain below its tensile strength even at elevated
temperatures (~1100°C) [kays93].
Adhesion of sensor films
Since the sensors used in high-temperature applications contain noble elements
such as platinum, it is generally difficult to achieve good adhesion to the
6
insulating film. Deposition in a slightly reactive atmosphere [gode87], and a high
energy sputter-deposition process [gran85] were proposed as solutions to this
problem.
Protective coatings
The use of protective coatings has been advocated by all groups whose work is
reviewed above. They are relatively thin (0.5-2.0µm) films of SiO2 or Al2O3 and
their purpose is to prolong the usability of the underlying sensors, especially in
the hot corrosive environments encountered in turbomachinery equipment.
However, these coatings are impractical when there is mechanical contact
between the tool or the component and the environment, as is the case in most
manufacturing or drilling tools. It is evident that, in these situations, a far
tougher coating is necessary and engineering alloys are the best candidate
materials. The only drawback is that such coatings have to be grown by high
temperature processes.
1.3 Basics of Sensor Physics
1.3.1 Strain Sensors
One-dimensional strain is the ratio of the incremental change of length to the
length of an object:
dl
ε = ---l
(1.1)
Its usefulness lies with the fact that it can be used to infer the stress applied to an
object and therefore indirectly define its stress state. Of course, this can only be
done accurately if there is a well known constitutive model that describes the
relation between the two quantities (stress and strain). For the family of materials
7
that behave elastically below a stress level this is relatively simple as the uniaxial
stress is related with the uniaxial strain with a linear formula:
σ = ε⋅E
(1.2)
where E is the Young’s modulus of the material, a material property giving a
measure of its stiffness and the strength of the interatomic bonding.
The most common method to measure strain makes use of the piezoresistive
effect, according to which the electrical resistance of a conductor changes when
the conductor is mechanically deformed. For example, let us assume a cylindrical
electrical conductor of length l, cross-sectional area a, and total volume V. The
resistance of this conductor is simply given by:
2
l
l
R = ρ ⋅ --- = ρ ⋅ ---V
a
(1.3)
where ρ is the resistivity or specific resistance, a material property of the
conductor and generally a polynomial function of temperature. When a uniaxial
tensile stress is applied to the conductor, the latter will be elongated, but, if the
stress is below the yield strength of the material, its volume will remain constant.
Differentiating Eq.(1.3) with respect to the length and rearranging the terms with
the aid of Eq.(1.1) yields a relation between the normalized incremental
resistance and the strain of the conductor:
dR
------ = S e ⋅ ε
R
(1.4)
where Se is the gauge factor or sensitivity of the strain sensor, and for most metals
ranges from 2 (ideal value) to 6 [frad93].
From Eq.(1.4) it is evident that the normalized change of the resistance of a
conductor that is subjected to a uniaxial stress can give information about the
8
strain in the conductor. The resistance change can be very accurately monitored if
the conductor is part of a balanced bridge circuit. Actually, this is the method
employed by strain indicator devices to report a strain value based on the
resistance change of the gauge [bray94].
The thin-film strain gauge is usually grown on an insulating foil that is
subsequently attached with a strong adhesive cement to the surface of the object
whose strain state is of interest. There is a complication arising from the fact that
the gage factor depends on the temperature of the conductor through its
resistivity and for this reason a special alloy (constantan) has been produced that
exhibits a very low temperature coefficient of resistivity (~10ppm/K) [buch91,
frad93]. The theoretical value of the coefficient is derived from a linear fit of the
polynomial law that governs the resistivity dependence on temperature.
Constantan is a copper-nickel alloy and its nominal composition is 50-65% Cu
with Ni in the balance [blat76]. It is widely used as a strain gage material in order
to minimize temperature effects on the resistance variation. The latter can be
completely accounted for with the use of a dummy gage. The dummy gage
should be on the same material and at the same temperature as the active gage
but it should not be loaded. Compensation of the thermal strain of the active
gage is achieved by connecting the two gages in adjacent arms of the bridge
circuit [bray94].
1.3.2 Thermocouple Sensors
In 1821 T.J. Seebeck discovered the homonymous thermoelectric effect which
manifests as a flowing current in a loop made of two dissimilar conductors when
the two junctions are at different temperatures. If one of the two junctions is
open, an electric potential (Seebeck emf) will appear between the two free
terminals. In this case the combination of the two conductors is referred to as a
thermocouple and such a device, when properly calibrated, can be used to
measure temperature.
9
The physical basis of this effect is the temperature dependence of the Fermi-Dirac
function. Using purely quantum-mechanical principles, Mott and Jones [mott58]
have derived expressions for the absolute thermoelectric power (or absolute
Seebeck coefficient) of noble metals and transition elements (above their
transition temperature). For reference purposes, the two expressions are
presented in Eq.(1.5) and Eq.(1.6), respectively [bedf90]:
2
2
–π k B T
S = --------- --------2 eE F
( µV ⁄ K )
(1.5)
2
2
kBT
–π
-------- -------------------------S =
6 e(E0 – EF )
( µV ⁄ K )
(1.6)
where EF is the Fermi energy, E0 is the highest energy level in the (incompletely
filled) d-band, kB is the Boltzmann constant and e is the electron charge.
However, Eq.(1.6) is not useful below the Curie temperature of the transition
elements or their magnetic alloys as the magnetic behavior of the electrons was
not considered in its derivation. A thorough discussion of the thermoelectric
transport properties of metals with reference to special problems in transition
elements can be found in [blat76].
In the case of a thermocouple the resulting thermoelectric potential will depend
on the difference of the absolute Seebeck coefficients of the conductors and the
temperature difference between the “hot” junction and the terminals (“cold”
junction) [poll85]:
dV ab = [ S a ( T ) – S b ( T ) ] ⋅ dT = S ab ( T ) ⋅ dT
(1.7)
where Sab is the differential Seebeck coefficient of the thermocouple, also referred
to as sensitivity of the thermocouple.
10
This coefficient is usually reported in µVK-1 or mVK-1, is approximated by a
linear function of temperature, and does not depend on the nature of the
junction. The most common type of thermocouple (type K) is formed by joining a
Ni0.9Cr0.1 conductor with a Ni0.94Mn0.3Al0.2Si conductor and has a sensitivity of
40.6 µVK-1 at 298K when the cold junction is held at 273K. Its sensitivity remains
reasonably constant for “hot” junction temperatures up to ~1300K. The alloys are
referred to with their respective trade names: chromel (or chromel P) and alumel.
Since the thermoelectric properties depend on the carrier transport mechanisms,
it is expected that imperfections and/or surface and grain boundaries in thin
films will affect their magnitude due to increased carrier scattering. In particular,
thin-film thermoelectric materials exhibit a decrease in their Seebeck coefficient
with decreasing thickness [balt94].
1.4 Deposition Processes
1.4.1 Sputtering
During the vapor deposition of a thin film we can identify a sequence of distinct
process phases: the generation of the species, its transport to the substrate
surface, the adhesion of the particles to the substrate, the nucleation of the film
and its growth. The term “sputtering” refers to the species generation, a process
that is described below.
It is possible to ionize, in a sustainable fashion, a mass of gas that is between two
electrodes by applying a voltage V across them. Then the free electrons, which
exist in very small concentrations in the gas, will be accelerated towards the
anode. In their trajectory they will interact with molecules of the gas or any other
evaporants, unless the distance d between the electrodes is so small as to allow
them to complete their travel uninterrupted. If the pressure P of the gas is low
11
enough, the electrons will have acquired enough kinetic energy as to ionize the
gas molecules, thereby injecting more free electrons in the space between the
electrodes. This cascading electron multiplication by gas-phase ionization is the
mechanism of gas breakdown in an electric field. Plasmas can be ignited and
sustained via this mechanism.
Apparently, the key parameters for this effect, for a certain gas species, are the
applied voltage V, the pressure P of the gas, and the distance d between the
electrodes. In fact, the pressure P is related to the mean free path le:
l e = ( σ n ⋅ n ) –1
(1.8)
where n is the molar density (analogous to P via the ideal gas law), and σn is the
collision cross-section, which depends on the type of collision and the electron
energy.
The generated gas cations are attracted electrostatically toward the cathode (or
“target”) and impinge on its surface. The conservation of momentum and energy
dictate that the kinetic energy Tm imparted to surface particles is:
4m i m t
T m = ------------------------2- E i = γ m E i
( mi + mt )
(1.9)
where mi and mt are the masses of the impinging and the target particles
respectively, and Ei is the energy of the impinging ion.
For particles of roughly equal mass the transfer of energy is very efficient. Only a
few tens of eV are enough to either displace a surface atom into the bulk of the
target or dislodge a nearby atom to the vapor phase with a few eV of initial
kinetic energy. The latter process, namely the erosion of the cathode surface due
to impingement of cations produced in a glow-discharge, is called sputtering and
12
is used to produce vapor of a solid element, alloy, or compound that will
condense on a substrate and form a thin film.
In its basic configuration, a sputter-deposition chamber contains a cathode to
which the “target” is attached. The target is usually a disk or rectangular-shaped
piece of the material to be deposited. If the target is conductive, the cathode is
connected to a dc power supply. In the case of an insulator the cathode is
capacitively coupled to a rf generator. Then, a high purity noble gas, generally
Ar, is injected in the chamber, which must have been evacuated to a pressure that
guarantees very low concentration of unwanted vapors.
The substrate is usually positioned directly across from the target at a distance of
6-10cm. It can be at a floating potential, grounded, or biased at a dc or rf voltage.
The application of a voltage to the cathode will instigate the breakdown of the
gas and lead to the formation of plasma. This setup is described as “diode”
sputtering.
We can confine the plasma near the target and sustain the discharge at lower
pressures and voltages by inserting magnets behind the target and thus
effectively extending the path of the ionizing electrons. This technique is referred
to as “magnetron” sputtering. Typical target voltages are in the few hundred
Volts range, and by virtue of Eq.(1.9), the sputtered particles retain a good
fraction of this energy. However, collisions with the noble gas atoms during their
flight to the substrate will diminish their initial ejection energy.
Microstructure
As the sputtered species are ejected with an initial kinetic energy some of them
are intercepted by the substrate. A fraction of them, dictated by the sticking
coefficient, will be adsorbed to the surface and form nuclei. Species that arrive
later will either form new nuclei or be added to pre-existing ones. When the
nuclei cover the whole surface, then nucleation ceases and the newly arriving
species contribute only to growth. The decisive factors for the final
13
microstructure is the temperature of the surface and to a lesser extent the energy
input to the substrate-film system by other methods, such as the bombardment
by neutralized species that bounce off the target surface, as well as the flux of the
incoming species.
Elevated substrate temperatures allow for the surface diffusion of the adsorbed
particles and thus promote a denser film. The diameter of the columnar grains
generally scales with the temperature of the substrate. Low impinging fluxes will
promote strongly textured films. When the substrate is at low temperatures,
surface diffusion is limited or even quenched, and the film grains have a fine
columnar structure which may exhibit voids caused by the self-shadowing effect
[smit95]. In general, sputtering at low working pressures (larger mean free paths,
fewer energy-dissipating collisions for the impinging species, lower angular
spread of incident angles) is thought to limit substantially the self shadowing
effects and produce finely grained yet dense films even at low temperatures.
Stress state
Sputtering can be responsible for what is called intrinsic or growth stress in the
deposited film. Of course, a sputter-deposited film can have stresses that are
caused by thermal strain mismatch (film and substrate having different thermal
expansion coefficients and elastic constants, and deposition occurring at high
temperatures) or by lattice parameter mismatch (“epitaxial” stress). It can also
exhibit a tensile stress due to atomic attraction between microclusters that are in
close proximity.
However, there can be a compressive stress component that arises due to the
bombardment of the film surface by energetic ions or neutrals. These projectiles
can either be implanted into the film and/or knock surface atoms into interstitial
positions and bring them into closer proximity. At low temperatures, the
displaced atoms can be “frozen” in the new positions and give rise to a
compressive stress. The case of momentum transfer that builds compressive
14
stress is very analogous to the shot-peening process used to compressively stress
the surfaces of bulk materials, so it is often called “atomic-peening”, or “ionpeening”. The “peening” effect is mostly visible at low working pressures, where
the mean free path is large enough to allow the deposited particle to hit the
substrate surface with a larger fraction of its initial momentum.
Adhesion
As explained before, the sputtered species are adsorbed to the substrate surface.
Their adhesion depends on factors like the cleanliness of the substrate and the
type of bonding that can be developed between the substrate and the incident
particles. It is quite possible that an unclean surface, contaminated with
monolayers of water or organic molecules, will not allow the chemisorption of
the deposited particles. Instead, these particles will only be physisorbed to the
substrate with weak dipolar bonds and the film adhesion will be poor. Therefore,
it is imperative that both the substrate is rendered atomically clean in vacuo
immediately prior to the deposition, and the sealing of the vacuum chamber is
very good. The top surface monolayers can be removed either by a physical
process (sputter-etching, substrate heating), or by a chemical one (e.g.: oxidation
of organic molecules in an oxygen plasma, reduction of oxides in a hydrogen
plasma).
With a clean surface the deposited particles can be chemisorbed to the underlying
atoms via the formation of new molecular orbitals. The adhesion strength is then
determined solely by the type of chemical bond that can exist between the two
substances. Evidently, metal particles can bond very well to a substrate made of
the same material. Also, some metals form strong oxides (Ti, Cr, Zr, Al) and they
can strongly bond to oxidized surfaces. This property is used to promote
adhesion between two otherwise weakly bonding materials (e.g.: Au on SiO2), by
depositing an interlayer of a “glue” metal (e.g.: Ti). Furthermore, sputterdeposition, being an energy-enhanced process, provides the impinging species
15
with enough energy content to activate interfacial bonding, thereby resulting in
better adhesion than evaporative deposition.
Composition
It is interesting that sputtered alloy targets will produce a deposit whose
composition can be very close to that of the bulk alloy. The atomic flux leaving a
target with two constituents a and b is:
Q k = S Yk ⋅ i
(1.10)
where SY is the sputtering yield, and i is the Ar-ion flux, whereas k denotes the
element. It can be assumed that for a fresh target the fractional coverage of its
surface by one of the constituents is ƒk. However, due to the different sputtering
yields, the ratio of the ejected fluxes (ƒa Qa/ƒbQb) will not be equal to ƒa /fb. As the
sputtering proceeds, the element with the higher sputtering yield will be
depleted from a thin volume at the surface of the target (its depth being diffusion
limited), and, consequently, the surface coverage of this element will decrease
until the ratio of the ejected fluxes is ƒa /fb. So, the composition of the ejected flux
will be the same with that of the bulk target. If some further assumptions are
met, (namely the equal sticking coefficients, the matrix independence of the
atomic sputtering yields, and the absence of ejected dimers), then the deposited
film will also have the same composition as the bulk target.
Reactive sputtering
The reactive sputtering technique is a variant of the sputter-deposition method
used to produce either doped or compound films by allowing an additional gas
species (other than the noble gas) in the deposition chamber.
The outcome of the deposition (doped film or compound) is determined by the
state of the target: in mode A, or “metallic” mode, the deposit is a metal film
doped with atoms from the reactive gas, while in mode B, or “compound”, or
16
“poisoned” mode, the deposit is a stoichiometric compound. The main
parameter that places the system in one of the two states is the flow of the
reactive gas (usually O2 or N2). The variation of the total system pressure with
the particular flow follows a hysteresis curve and the existence of such a curve is
indicative of a reactive system. A generic hysteresis curve is shown in Figure 1.1.
Total pressure
Pn
P
PT
P’T
PAr
Q’T
QT
Reactive gas flow
FIGURE 1.1: Generic hysteresis curve in a reactive sputtering system.
Assuming that the sputtering gas (usually Ar) flow and the pumping speed
remain constant, there is a history effect in the variation of the system pressure
with the reactive gas flow. Initially, at zero reactive flow, the deposit is purely
metallic (if the target is metallic) and the total pressure is equal to the argon
pressure. If there were no discharge, the introduction of additional mass flow
into the system would cause the system pressure to follow the diagonal dashed
line (Pn(Q)). However, in the case of a discharge, and for small reactive gas flow
values, the total pressure remains essentially constant. The reason for this
phenomenon is the trapping or getter-pumping of the reactive gas in the
deposited film and the final result is a doped metallic film. The target is in the
“metallic” mode A.
17
When the reactive gas flow exceeds a value QT, then the pressure rapidly
increases to PT, which differs from Pn(QT) by an amount equal to ∆P. Higher flow
values will lead the system to higher pressures (P) but the difference
∆P(Q)=Pn(Q)-P(Q), for Q>QT, will be constant. This pressure difference
corresponds to the amount of the reactive gas that has been incorporated into the
deposit, which is now a stoichiometric compound. The target forms a thin
compound film on its surface and transitions to the “poisoned” mode B. The
target will remain in the “poisoned” state even for Q<QT and it will revert to the
metallic mode for flows below a critical value Q’T. At this state the vast majority
of the species ejected by sputtering of the target are compound molecules.
The previously described hysteresis is explained by the difference in the
sputtering yields between the metallic element(s) of the target and the compound
formed in the presence of the reactive species. Generally, the compound will
have a lower sputtering yield than the fresh, purely metallic surface and the
deposition rate will drop. Consequently, the rate at which the reactive gas is
getter-pumped in the deposit and consumed at free sites at the target surface will
drop too, and there will be an excess amount of the reactive gas which will
register as an increase in the total system pressure. The high transition flow QT
represents the state where the rate of compound formation on the surface target
(or “target poisoning”) becomes higher that the rate at which the target is
sputter-cleaned of the compound. This trend, once the target becomes
“poisoned”, is reversed only at the lower transition flow value Q’T.
Due to the difference in the sputtering yields it is evident that the compound and
the (doped) metal will have different deposition rates. This difference becomes
dramatic when the compound is an insulator, e.g.: Al2O3. However, there is an
operating window of reactive flow values that makes reactive deposition of
insulators attractive from a deposition rate point of view. (It is obvious that
reactive sputtering is more versatile than rf-sputtering as it is capable of
18
depositing different compounds from a single target, e.g.: AlN and Al2O3.) If the
reactive gas flow is kept within a narrow range around QT, then it is possible to
deposit a compound film with good stoichiometry at a rate that is comparable to
that of the metallic film.
The difficulty associated with this procedure is keeping the total pressure at a
level between PT and PAr, while the flow is kept at a value of QT. When the
reactive gas flow is equal to the transition value QT, the system is in an inherently
unstable equilibrium: a small perturbation of the flow towards higher values will
increase the “target poisoning” rate, so less reactive gas will be consumed, thus
the reactive gas partial pressure will increase and this will lead to a further
increase of the “poisoning” rate until the equilibrium pressure of P1 is finally
reached. At this pressure the target is fully “poisoned” and the deposition rate is
very low.
Alternatively, a perturbation towards lower flow values will result in faster
sputter-cleaning of the target, faster consumption of the reactive gas, and
lowering of the system pressure to the equilibrium value of PAr. At this stage the
target is in the metallic mode, the deposition rate is high, but the deposit is a
doped metallic film and not a stoichiometric compound.
From the discussion above it can be inferred that the most important parameter
in reactive sputtering is the flow level of the reactive gas. Its control is essential
for the high-rate deposition of stoichiometric compounds due to the unstable
equilibrium that is characteristic of this state.
Pulsed-dc sputtering
In the case of reactive sputter-deposition of insulating films from a metallic target
(e.g.: Al2O3) the insulating species that form on the surface of the metallic target
and get deposited over the anode-ground alter the electrical characteristics of the
discharge, building effectively a capacitor between the plasma and the
19
electrodes. In particular, the growth of an insulating layer that covers gradually
larger parts of the anode or ground (which is essentially the chamber walls)
makes it more difficult for the plasma electrons to find a conducting surface and
close the circuit. The result of the “disappearing anode” is extensive arcing –
which damages the deposited film –, stray plasma, and long-term plasma
instability [schi93]. The problem can be addressed for the mid-term with
“concealed” or rotating anodes but it does not disappear altogether.
Pulsed-dc sputtering is manifested by superimposing a low frequency square
waveform on the dc signal that dives the magnetron target. The superimposed
waveform can be unipolar or bipolar. In the latter case it can be symmetric or
asymmetric. If a single target is used and the bipolar mode is selected in order to
enhance the removal rate of the insulating species from the “poisoned” target
surface, then the asymmetric waveform is preferred in order to avoid sputtering
of the anode (chamber walls) and consequent contamination of the deposit.
The operating principle behind bipolar asymmetric pulsed-dc sputtering lies
with the enhanced momentum of the argon ions hitting the target when the
voltage returns to its negative plateau (point A in Figure 1.2).
+100V
time
-300V
A
FIGURE 1.2: Asymmetric bipolar dc voltage as applied to a magnetron target.
At that point the parasitic capacitor, which has formed by the insulating film
grown on the target surface, has been fully charged and reached a potential equal
20
to -100V (as the metallic target was held at 100V). As the voltage reversal
happens and the metallic target returns to -300V the effective voltage that
accelerates the argon ions to the cathode is -400V, and therefore the sputtering
yield increases, effectively cleaning the target from the insulating species [sell96].
The frequency of the pulse is generally in the 50-250kHz range and the pulse
width, which governs the duty cycle, is in the 500-2000ns range.
The use of pulsed-dc sources for reactive magnetron sputtering has allowed the
stable and lengthy deposition of insulators at rates comparable (60%-70%) to
those for the pure metals. The latter rates are significantly higher (by at least an
order of magnitude) than the rates for insulators from compound targets
achieved with rf-sputtering.
1.4.2 Electroplating
Electroplating is the process of producing a coating by electrolysis. According to
Faraday’s laws, when one Faraday of electricity (~96490As) is transmitted
between two electrodes immersed in an electrolytic solution (“bath”), 1 g-equiv
(gram-equivalent) of metal will be deposited on the cathode.The mechanism by
which this happens is described below.
When a piece of metal is immersed in a aqueous solution containing its ions
some of the atoms will leave the lattice and dissolve into the solution, becoming
hydrated. At the same time, some of the ions will leave the solution and attach to
the electrode. The potential of an electrode at which the two opposite reactions
have equal rates is called equilibrium potential E and depends on the metal, the
solution and its temperature, per the Nernst equation [lyon74]:
0
E = E M n + + RT ⁄ ( nF ln a M n + )
(1.11)
21
where E0Mn+ is the standard potential for reduction of the simple metal ion Mn+
to the atom M, R is the gas constant, T is the absolute temperature and aMn+ is
the activity of the metal ion in the solution.
If we consider two copper rods dipped in a water solution of copper sulfate, and
we connect them to an external direct current source, we will raise the potential
of the rod connected to the anode and lower the potential of the rod connected to
the cathode by supplying more electrons to the latter. The fact that the cathode
assumes a potential lower than the equilibrium potential means that more copper
ions will leave the solution and attach to that electrode, thereby closing the
electric circuit and increasing the mass of the cathode. The reverse will happen in
the anode. This is a very basic electroplating “cell”.
Microstructure
The electrodeposits are generally microcrystalline in nature. Crystal grains
nucleate more easily at kinks, edges, or steps on the lattice of the substrate and,
initially, grow laterally in a monolayer fashion as the ad-ions diffuse across the
surface. When the grains encounter adsorbed impurities, or other growing
grains, they start growing outward. At low currents and very low concentration
of impurities it is possible to have “epitaxially” dictated crystallographic
directions in the growing grains, assuming, of course, that the lattices of the basis
metal and the coating are similar. It is also possible to affect the growth pattern
by providing enhanced convection of fresh ionic species to the vicinity of the
cathode, e.g.: by stirring.
Stress state
Stresses in the electrodeposits arise from lattice mismatch at the basis-coating
interface, from coalescence of adjacent crystal grains and, more commonly, from
incorporation of foreign substances in the lattice. Such inclusions as oxides,
hydroxides, water, sulfur, carbon, hydrogen, and metallic impurities (foreign ions
22
in the bath) distort the lattice and produce stress fields around them which may
manifest as macroscopic stresses [read74].
Adhesion
As the first monolayer of the elctrodeposit is grown on the surface of the basis
metal, its atoms engage the lattice forces of the latter. In the absence of impurities
from the basis-coating interface the bonding is very strong and approaches the
strength of the bonds among the basis metal atoms. Only improper substrate
preparation, or pronounced lattice mismatch between the basis and the coating
materials, can be the cause for poor adhesion [lyon74].
1.4.3 Laser Assisted Deposition
A number of layered manufacturing techniques make use of a high-power laser
beam to create fully dense metal parts by welding new layers onto substrates.
The beam is focused on the surface of the substrate and thus creates a pool of
molten material. Then, more material is injected into the pool in powder form
and the beam is moved relative to the substrate. The powder feeder follows the
laser motion and continuously supplies more material to the molten metal pool
which in turn solidifies and fuses to the substrate as the laser moves on its
predefined trajectory. An alternative method of supplying new material is to use
a pre-deposited powder bed on the substrate.
Particular incarnations of this additive welding-base technique include Laser
Engineered Net Shaping (LENS) [beam97], Directed Light Fabrication (DLF)
[beam97], Control Metal Build-Up (CMBU) [beam97], LaserCast [hous97], Laser
Direct Casting [mcle97], and Laser Deposition in Shape Deposition
Manufacturing (SDM) [nick99, link99].
In SDM, a sacrificial support material is used in order to allow the realization of
overhanging structures and pre-assembled mechanisms, and each layer is
planarized and shaped after its deposition [merz94].
23
1.5 Dissertation Outline
The subject of the work presented in this dissertation is the embedding of
thermo-mechanical sensors inside metallic components, tools, and structures.
The main focus of the research has been the development of the processes
necessary to achieve both the fabrication of the sensors and their subsequent
enclosure (embedding). Chapter 2 describes the scheme we propose in order to
produce the sensors and the layers necessary for their protection and embedding.
The full details of the fabrication procedure are described in Chapter 3. The
characterization results of the various components – with respect to their most
relevant properties – are presented in Chapter 4. The results of a heat transfer
finite element analysis, which was employed to predict the temperature at critical
regions of the structure, are the subject of Chapter 5. The final chapter contains
the summarized conclusions for the research described in the dissertation and a
few guidelines regarding the implementation of the technology developed at the
Rapid Prototyping Laboratory at Stanford University towards the production of
an instrumented tool.
24
2
Proposed Solution
In this chapter a solution to the embedding of thermo-mechanical sensors inside
metallic components, tools, and structures is proposed vis-a-vis the challenges
that were outlined in the previous chapter.
The major obstacles towards the fabrication of a tool with embedded sensors can
be identified as the electrical insulation of the sensors and their protection from
the high temperature processes. The problem of compatibility between the
various processes becomes more evident during a change of the production
environments. Particularly, it is a matter of cleanliness and surface quality when
a part is transferred from large-scale manufacturing to thin-film processing and a
matter of physical (mainly thermal) and chemical protection during the reverse
procedure.
Bearing the above in mind, our proposed solution can be outlined as follows:
2.1 Cleanliness and Surface Quality
Before entering the thin-film processing stages the semi-completed part will have
to be thoroughly de-greased, cleaned and dried. Its surface that will serve as a
substrate for the thin-film sensors will have to be polished and devoid of any
residual chemicals. It is important that it is sputter-etched inside a vacuum
25
CHAPTER 2: PROPOSED SOLUTION
chamber to ensure that the contaminated top monolayers have been removed
before any deposition takes place.
2.2 Insulation
The insulating material of choice should be easily deposited via a thin-film
method and be able to withstand elevated temperatures during subsequent
processing and operation of the component. Possible candidates are the oxides of
aluminum, silicon, and tantalum, and the nitride of silicon. All of them can be
deposited by rf-sputtering of compound targets or reactive sputtering of the
elemental targets in an atmosphere that contains oxygen or nitrogen.
Obviously the insulating layers should fully enclose the sensors except for the
contact pads which must remain exposed – but insulated from the substrate. In
the case of the temperature sensors it also makes sense to have the junction at
direct contact with the surrounding matrix.
2.3 Sensors
The sensors can be deposited by sputtering targets of materials used for the
fabrication of large-scale sensors, e.g.: constantan, copper, alumel, chromel, etc.
Sputtering has the intrinsic advantage of achieving a deposited film that has the
same composition of the bulk target material assuming that all the molecules
have roughly the same sticking coefficient.
The patterning of the sensors can be achieved by means of photolithography and
micromachining (e.g.: lift-off, etching, ion etching).
26
CHAPTER 2: PROPOSED SOLUTION
2.4 Thermal Protection
Our goal is to protect the thin films of the insulation and the sensor during the
completion of the component via high-temperature deposition or diffusion
processes. This can be facilitated by providing an alternative fast path to the heat
input during the high-temperature deposition so as to be conducted mainly to
the thick metallic substrate (of higher heat capacity) than to the sensitive thinfilm multilayer.
This path can be produced by growing a thick layer of a material with high
thermal diffusivity, such as copper. This layer, which has to be grown by a lowtemperature technique (e.g.: electroplating), will extend over and beyond the
insulation layers as to have immediate contact with the metallic substrate. Of
course, it will have to be insulated from the sensors by the top insulating film.
2.5 Embedding
The embedding of the sensors and the protective layers in the metallic structures
can be achieved by a high-temperature deposition technique akin to those used
in layered manufacturing. For example, the fusion of a powder bed to the
underlying structure can be achieved with a high-power laser. Alternatively, a
more benign plasma-spraying process can be used if the full density and
chemical bonding achieved by cast-like processes are not necessary. Diffusion
bonding might also offer an additional option. However, a process that can lend
itself to freeform fabrication is always preferable, as it can achieve the shaping of
the added material into useful geometries.
27
3
Fabrication
The procedures and processes necessary to fabricate a thin-film thermomechanical sensor and embed it in a metallic matrix are the focus of this chapter.
Even though the processes described guarantee the successful production of such
a device, it is expected that there can be optimizations which, when applied upon
those processes, will result in a higher yield and/or faster production time.
3.1 Substrate preparation
Ideally, the thermo-mechanical sensors should be deposited upon an internal
surface of a tool or other structure during its fabrication with a layered
manufacturing technique of such a tool with the SDM process [“Laser Assisted
Deposition” on page 23]. However, for practical reasons of availability, substrates
that could simulate the part surface in terms of materials properties were used.
3.1.1 Cutting
The standard substrates used in the process are 3mm-thick stainless steel pieces
with nominal lateral dimensions of 50x50 or 50x25mm. These pieces are cut
from square 30x30cm plates using band saws and cutting wheels. As a result the
exact dimensions of the substrates are systematically smaller than their nominal
28
CHAPTER 3: FABRICATION
values by up to 3mm.The stainless steel is of the 304L grade with a nominal
composition [asm94] that is described in Table 3.1.
TABLE 3.1: Nominal composition of 304L stainless steel.
Cr
Ni
Mn
C
Fe
18-20%
8-12%
2%
0.03%
bal.
The reasons behind the selection of the particular material were the wide
availability, machinability, and mostly the corrosion-resistant character, which
would allow its use in a variety of processing environments.
3.1.2 Grinding and polishing
The substrates need to be ground and polished for two reasons. First, a smooth
surface can ensure easier and more efficient cleaning and provide little chance for
large size particles to mechanically attach to it. Moreover, a smooth substrate is
free of cusps which may function as stress concentration regions for the ceramic
film that will be deposited above them [stor96, nisk98].
The grinding process involves 5 grades of grinding paper with increasingly finer
SiC particles denoted by grit numbers: #320, #400, #600, #800/2400, #1200/4000.
Usually, 4 minutes of grinding over every sandpaper are sufficient for good
results when the paper is fresh and the substrate orientation is rotated by 90
degrees after 2 minutes. A rotation speed of 180-220rpm seems to work best and
the flushing of the substrate with water between successive sandpapers is
absolutely necessary to avoid cross-contamination with lower grit particles. It is
good practice to use latex gloves for handling the substrates throughout the
whole preparation stage.
The polishing process involves 3 different suspensions of alumina particles in
distilled water. The sizes of the particles used are 1.0, 0.3, and 0.05µm. The
optimal time for polishing has not been established, however, the disappearance
29
of scratches that were left by the previous polishing (or grinding) step is
adequate evidence that one may proceed with the next suspension. However, a
bluish hue covering part of the substrate was observed after prolonged polishing.
The rotation speed used is in the 300-500rpm range. It is always beneficial to use
fresh polishing cloth and it is mandatory to flush the substrate thoroughly with
distilled water between polishing steps. The average roughness of the surface
achieved by this method is always smaller than 20nm, as measured by a Tencor
Alpha-Step stylus profilometer for scan distances in the 80-2000µm range.
In the quest for a polished surface, free of attached particles, the substrates were
electropolished in a standard stainless steel polishing bath with a current of 5070A at a voltage setting of 8V dc. The results did not compare well with those
achieved by mechanical polishing, as the surface revealed pits after 3 minutes of
electropolishing.
3.1.3 Cleaning
The substrates have to be cleaned after the polishing process. The sequence of
steps to that end is described in detail below and takes place inside a class-1000
clean room. In an effort to remove salts, oils and grease we rinse the substrates
sequentially in borothene, acetone and isopropanol. Then the substrate is blown
dry with dry nitrogen, rinsed with de-ionized water and dried again. The above
steps are repeated once more and then the substrates are immersed in a 1:15
solution of micro-detergent in de-ionized water and are placed in an ultrasonic
bath for 5 minutes. Finally the substrates are rinsed with copious amounts of deionized water and blown dry with dry nitrogen. It is expected that at least a
monolayer of the surfactants in the micro-detergent will remain on the surface of
the substrates [mbec98, also: “Discussion of cleaning processes” on page 33], as
well as carbon from the organic solvents.
In order to eliminate persistently adhered particles to the stainless steel surfaces
and enhance their passivation it was suggested to immerse the substrates in a
30
1:10 solution of nitric acid (70%) in de-ionized water for a few minutes and then
flush with de-ionized water [leve98]. The results of our attempts, in terms of
insulation failures due to pinholes, were not conclusive as to the effectiveness of
this method.
Other researchers [nasa86] have employed exposure of the substrates in UV
radiation in an ozone flow for 60 minutes to dissociate and remove organic
substances from the surfaces. However, they did that in a clean room
environment. This procedure was not replicated, as one of the goals of this work
is to create a process sequence that can be carried out with successful results in a
more traditional manufacturing environment.
Since the thin-film processing does not take place in a clean-room area, it is
necessary to take precautions that minimize the exposure of the substrates to the
environment outside the clean room during their transfer. For this reason it is
recommended to mount the substrates onto the substrate holder in the clean
room and transport the holder in a ziploc plastic bag filled with dry nitrogen.
3.1.4 Sputter-etching
The final step of the substrate preparation takes place inside the vacuum
chamber just prior to the deposition of the insulation layer. Its purpose is to
remove any electrostatically attached particles or chemisorbed substances form
the substrate surface by etching a thin layer off the top of the substrate material.
This process uses accelerated argon ions to sputter material from the substrate
instead of the target and is called sputter-etching, sputter-cleaning or backsputtering. The chamber is evacuated to a base pressure better than 10-4 Pa, then
argon is injected and the pump is throttled to maintain a pressure of 6Pa. A
pulsed-dc signal is applied to the substrate holder at a constant power setting of
100 W. The full details of this process are listed in Table 3.2.
31
TABLE 3.2: Sputter-etching process parameters.
Base pressure
<10-4 Pa
Argon flow
150±3sccm
Argon pressure
6.0±0.2Pa
Power (controlled)
100 W
Pulse frequency/width
201kHz/1056 ns
Discharge voltage
325-330V
Etching time
600s
Since there is no magnet behind the substrate stage, it is necessary to use a high
pressure of argon to ensure there are enough ionization events and consequently
a self-sustaining electric discharge through the rarefied gas. We also have to use a
shutter between the substrate and any facing targets to avoid contamination of
the target with material sputtered from the substrates.
A measurement for the sputter-etching rate of the 304L stainless steel substrate
was taken by step profilometry: a piece of silicon wafer was placed on top of the
steel surface and after 600s of sputter-etching the height of the created step was
measured. The resulting rate was estimated at 3nm/min. However, this rate
results in just 30nm of height difference. Bearing in mind that the average
roughness of the polished steel surface is of similar magnitude (10-20nm) and
that the step was not very sharply defined, it might be necessary to create such a
step with different methods and longer etching times.
In the pulsed-dc mode, the rate of 3nm/min is achieved by sputter-etching
during only 79% of the duty cycle. During the rest 21% the substrate assumes a
positive voltage and no sputter-etching occurs. Ideally, a purely dc signal of the
same power level would result in a rate of 3.8nm/min. The reason for using
pulsed-dc mode is that it can be successfully used for conducting and insulating
substrates alike and, more importantly, it can sustain a discharge for longer times
32
at lower pressures than the pure dc mode, which is very prone to arcing [cf.
“Pulsed-dc sputtering” on page 19].
3.1.5 Discussion of cleaning processes
In order to evaluate the relative effectiveness of the different cleaning steps, three
polished 304L stainless steel substrates were prepared, referenced as CLD, CLS,
and CLN, and were cleaned in three different manners. Substrate CLD was
cleaned in a micro-detergent and de-ionized water solution (1:15) inside an
ultrasonic bath for 5 minutes. Substrate CLS was cleaned repeatedly with the
organic solvents mentioned in “Cleaning” on page 30, and substrate CLN was
immersed in a 7% v/v nitric acid solution. All of them were subsequently
thoroughly rinsed with de-ionized water and blown dry with nitrogen.
All substrates were scanned in a XPS system for a rough compositional analysis.
Then, all samples were sputter-etched for 90s and re-scanned to gather
compositional data. The surfaces were sputter-etched for another 210s before the
final scan. The rationale of this process sequence was to observe any changes in
the surface composition – and particularly carbon, oxygen, and iron contents –
with depth. The results are summarized in Table 3.3 and Table 3.4 .
TABLE 3.3: Surface composition of 304L substrates cleaned with different methods.
Substrate
Carbon
(at%)
Oxygen
(at%)
Iron
(at%)
Chromium
(at%)
CLD
23.4
57.3
8.1
8.5
CLS
21.0
60.5
9.5
6.3
CLN
28.7
53.0
8.2
8.2
It can be seen than in all cases, oxygen and carbon constitute 81-82% of the
as-prepared surface layers, irrespective of the cleaning process that was followed.
The origin of the oxygen is mainly the water molecules that are attached to the
substrate surface by Wan der Waals forces. We believe that this contamination is
33
TABLE 3.4: Atomic content of O, Fe, and Cr in 304L surfaces after sputter-etching.
Substrate
CLD
CLS
CLN
Etching time
(s)
Oxygen
(at%)
Iron
(at%)
Chromium
(at%)
90
13.5
51.7
13.8
300
9.9
56.1
14.3
90
14.8
59.2
10.1
300
11.3
59.5
10.2
90
10.1
60.7
11.2
300
8.4
64.2
16.1
due to the exposure of the samples to the atmospheric humidity during transport
and/or storage. The carbon most probably originates from the organic solvents
residue, and the organic agents in the micro-detergent or the atmosphere. In any
case, it is remarkable that the treatment with organic solvents leaves the least
carbon contamination on the surface.After 90s of etching there is no detectable
carbon in the substrates and this is evidence that the carbon contamination is
strictly a surface phenomenon and can be completely removed even by minimal
sputter etching. The significant oxygen content in the first few nanometers
beneath the topmost surface layers is probably due to diffusion of atomic oxygen
and bonding with active elements such as iron.
3.2 Bottom insulation layer
The first layer to be deposited will have to provide the sensors, and especially the
strain gages, with adequate electrical insulation from the stainless steel substrate.
Typical candidates for this function are the oxides of silicon and aluminum as
well as silicon nitride and tantalum pentoxide. In principle, and practice, all of
the above can be deposited in thin film form via rf-sputtering from a compound
target or reactive sputtering from a metallic target in an atmosphere that contains
oxygen or nitrogen. In our case, reactive sputtering was the only available
solution due to the lack of a rf generator and matching system and also due to
34
the high rates that can be achieved with reactive deposition. The high rates are
necessary in cases with poor particle control, since thicker films will be less likely
to suffer from pinhole problems caused by particle contamination.The price to be
paid for choosing a reactive system is the difficulty in the control of the
deposition process.
We chose aluminum oxide as the preferred insulating film for reasons based on
thermal expansion coefficient compatibility, and target cost. These parameters are
listed for the aforementioned insulating films in Table 3.5 below:
TABLE 3.5: Material properties for insulating thin films.a
Material
Dielectric strength
(kV/mm)
Thermal expansion
coefficientb
(µ K-1)
Cost of metallic
target
(US $cm-3)
Al2O3
9.9-15.8
8.0
~12
SiO2
15-25
0.3-0.4
~20
Ta2O5
>200c
Si3N4
15.8-19.8
~40
2.5-3.5
~20
a. [buch91], unless noted otherwise
b. Mean thermal expansion coefficient of 304L stainless steel, in the 20-400°C range:
18 µε K-1 [asm94].
c. [chen97]
The concept of reactive sputtering has been presented in “Reactive sputtering”
on page 16. Hereafter, the focus is turned on the issues of the process as appeared
in our system.
Our goal was to produce an aluminum oxide film that did not suffer from
pinhole problems, was as close to the ideal stoichiometry as possible, and could
be deposited in a reasonable time (1-2hr). The solution to these issues would
entail certain compromises, since in order to guarantee perfect stoichiometry the
deposition rates would have to be minimized and a thick film (in the order of a
few µm) would take several hours to grow. The key to the solution is the
35
determination of the transition oxygen flow (or oxygen content) that would
permit the metallic target to remain partially “poisoned”, therefore allowing high
rates of deposition with a stoichiometry very close to the ideal.
3.2.1 Determination of transition oxygen flow
In order to establish this optimum concentration of oxygen in the sputtering gas
mixture, the levels of the discharge voltage and the total pressure during the
sputtering of an aluminum target were recorded versus a gradually varying
oxygen flow. The argon flow was kept constant at 80sccm and the discharge
current was preset at 0.80A. The power supply was switched to the pulse mode
with a frequency of 181kHz and a pulse width of 1056ns. The goal was to
reconstruc of the hysteresis curves for both the voltage and the pressure and thus
determine the optimum oxygen flow.
Two experiments (Figure 3.1) were conducted with targets in different stages of
their useful life. The first one, referenced as 980425, was a a relatively fresh target
which had been previously sputtered in an argon-oxygen atmosphere for a total
of 3 hours. The second target, referenced as 990416, had been used for 16 hours
prior to the experiment and was approaching the end of its life. For both
experiments the base pressure was less than 3.5x10-4 mPa and the argon flow was
kept constant at 80sccm. The turbomolecular pump was throttled so as to
maintain a pressure of 0.80Pa at 0.80A and zero oxygen flow (total pressure
equal to argon partial pressure).
In the case of target 980425 oxygen was introduced in the chamber in steps of
2sccm every 200 seconds, allowing for the stabilization of the voltage. The
stabilized voltage and the total pressure were then recorded. It must be noted
that the voltage is changing rapidly when the flow is close to the transition point.
Furthermore, the measured mass flow was consistently smaller than the set flow.
36
This discrepancy reached a maximum of 0.7sccm at the maximum set flow level
of 12sccm.
For target 990416 the oxygen flow was increased in steps of 0.2sccm every 150
seconds and the voltage and pressure were recorded in 50 second intervals. The
range of the set flow values was from 0 to 11.1sccm, which corresponds to a
measured range of 0-10sccm. Thus, the maximum discrepancy between set and
measured mass flow was 1.1sccm and occurred at the maximum set flow of
11.1sccm.
990416: old target
980425: fresh target
0.95
Voltage (increasing flow)
Voltage (increasing flow)
280
Voltage (decreasing flow)
0.92
Voltage (decreasing flow)
P = m1 + m2 * Qox
m1
m2
R
0.7778
0.0137
0.9995
0.85
200
240
150
0
2
4
6
8
10
0.75
12
Oxygen Flow (sccm)
m1
0.7830
m2
0.0137
R
200
0.9921
Pressure (increasing flow)
Pressure (decreasing flow)
Linear fit for pressure
0.80
Pressure (increasing flow)
Pressure (decreasing flow)
Linear fit for pressure
0.88
P = m1 + m2 * Qox
0.84
Pressure (Pa)
250
Voltage (V)
0.90
Pressure (Pa)
Voltage (V)
300
0.8
160
0
2
4
6
8
10
12
Oxygen flow (sccm)
FIGURE 3.1: Discharge voltage and total pressure vs. oxygen flow for two Al targets.
Argon flow: 80sccm. Current: 0.80 A. Partial pressure of argon: 0.80Pa
The marked reduction of the discharge voltage at high oxygen flows is caused by
the difference in the coefficients for emission of secondary electrons by ion
bombardment between the metallic and the oxidized aluminum. The Al2O3 thin
film, which is formed on the surface of the target at high oxygen flows, has a
much higher emission coefficient than metallic Al [mani80a].
Discussion of the high oxygen flow regime
One common observation for both cases is the linear dependence of the total
pressure on the oxygen flow at high flow values. The slope coefficient of the
linear fit for oxygen flows higher than 7sccm is 0.0137Pa/sccm. If we plot the
37
total pressure as a function of the oxygen flow when there is no plasma discharge
(Figure 3.2), and do a linear fit for flows higher than 7sccm, we get a slope
coefficient equal to 0.0138±0.0004Pa/sccm. This fact is in accordance with the
literature results [mani80b] where the pressure varies linearly with the flow of
the reactive gas when the target is fully oxidized, an effect that manifests itself at
high reactive gas flows.
Pressure (Pa)
0.95
0.90
0.85
P = m1 + m2 * Qox
0.80
0
1
2
3
4
5
6
7
8
Value
Error
m1
m2
0.7856
0.0138
0.0038
0.0004
Chisq
R
0
0.9975
NA
NA
9
10
11
12
Oxygen flow (sccm)
FIGURE 3.2: Variation of total pressure with oxygen flow.
Argon flow: 80sccm. Current: 0 A. Partial pressure of argon: 0.80Pa. The linear fit is
calculated for the flow values used in the linear fits in Figure 3.1.
In these cases the extra amount of oxygen introduced is not getter-pumped at
active aluminum surfaces in order to create more aluminum oxide. It only adds
to the total pressure. Similarly, when there is no discharge at all, any amount of
oxygen introduced to the chamber increases the total pressure.
The vertical offset of the linear fit curves is an indication of the amount of oxygen
being consumed while the target is in the fully “poisoned” state. Assuming
constant argon flow, discharge current, and pumping speed, the vertical offset
should scale inversely with the amount of oxygen that is consumed in the
process. The slope should remain the same since, in the fully “poisoned” state, it
only depends on the pumping speed.
38
Ideally, a constant pumping speed over the range of total flow values (sum of
oxygen and argon flow) would result in a vertical offset of 0.80Pa for the nodischarge case. However, the pumping speed is not constant at this pressure
range [leyb94]. It decreases as the total flow and pressure are increased, leading
to steeper slopes for the pressure-flow curves at high flow values. Consequently,
when the linear fits for high values of oxygen flow are extrapolated to zero
oxygen flow, the resulting offset is less than the actual pressure of 0.80Pa (see
Figure 3.2).
In Table 3.6 we summarize the results of the linear fit curves and juxtapose them
to the discharge voltage in order to show that our results are in agreement with
the expected behavior of the process. In particular, the offset without discharge
has indeed the highest value and decreases with increasing voltage (and power)
levels. The reason is that at higher power levels the target is being eroded at
higher rates which result in higher consumption rates of the available oxygen
and, therefore, in a lower partial pressure for oxygen.
TABLE 3.6: Linear fit parameters for high oxygen flows (Figure 3.1 and Figure 3.2).
Target
Offset
(Pa)
Slope
(Pa/sccm)
Voltage
(V)
980425
0.7778±0.0022
0.0137±0.0002
194
990416
0.7830±0.0037
0.0137±0.0005
164
N/A
0.7856±0.0038
0.0138±0.0004
0
Discussion of the history dependence of the transition point
Another interesting fact is the dependence of the voltage-flow curves on the
history of the targets. In Table 3.7 we summarize the most important attributes of
the two curves which are depicted overlaid in Figure 3.3.
It is evident that as the a target is used for repeated depositions, the hysteresis
loop becomes shorter, wider and more square and is displaced to the lower flow
39
TABLE 3.7: Characteristics of the voltage-flow hysteresis curves in Figure 3.1.
Target
Hours of
use
Vmax-Vmin
(Volts)
Vinita
(Volts)
980425
3
100
286
990416
16
77
240
a.
b.
c.
d.
∆Q at
Q’c
(sccm)
QTd
(sccm)
0.9
9.4
6.6, 5.8
1.2
4.8
4.6, 3.4
(Vinit-Vfin)/2b
(sccm)
Initial voltage at zero oxygen flow.
Width of hysteresis loop halfway between the initial and final voltage values.
Highest oxygen flow at which hysteresis is still observed.
Transition flow: value at which the voltage is Vinit-30 V.
350
16 hr target (incr. flow)
16 hr target (decr. flow)
3 hr target (incr. flow)
3 hr target (decr.flow)
Voltage (V)
300
Transition
point
250
200
150
Transition
point
0
2
4
6
8
10
12
Oxygen flow (sccm)
FIGURE 3.3: Transition voltage and flow for two Al targets (Figure 3.1 & Table 3.7).
values. The flow value Q’ is essentially the flow required to bring the target to the
fully “poisoned” state.
The most important parameters are the initial voltage Vinit and the “transition
flow” QT. Prior to every deposition the target is sputtered for 200-300 seconds in
a pure argon atmosphere (zero oxygen flow) and the measured voltage of the
discharge is Vinit. The “transition flow” QT, and particularly the first value which
corresponds to an increasing flow, has been selected as the operating oxygen flow
level. It results in a partially “poisoned” target and the associated voltage drop
40
can be maintained at that level over prolonged times with minor adjustments in
the flow. Experience has shown that higher oxygen flows will result in more
dramatic and even unstable voltage drops, while lower flow values are also
prone to unstable voltage changes, albeit in the opposite direction. Furthermore,
with oxygen flow at the QT level, the deposit has a stoichiometry close to the
ideal and the deposition rates (0.4-0.7nm/s) are comparable to the ones observed
for metals.
An interesting result of the above is the drift of the transition point towards
lower oxygen flows and lower voltage and power values as the targets get older.
We have tracked the QT and power values for a series of deposition sequences for
one target and we present the results in Figure 3.4. The most likely cause for this
drift is the oxidized condition of the target at the end of each deposition
sequence. It can be speculated that after every reactive sputtering process the
target becomes more heavily poisoned, and therefore less oxygen is needed in
order to bring the target in the transition state.
(a)
(b)
240
8.0
6.5
Power
220
6.0
200
6.5
180
6.0
5.5
160
O 2 flow (sccm)
7.0
Power (W)
Transition O2 flow (sccm)
7.5
5.5
5.0
Q = a + b*S
T
5.0
140
4.5
O flow
4.0
2
0
1
2
3
4
5
6
Deposition #
7
8
9
120
10
Value
Error
a
b
-2.2602
0.0383
0.5210
0.0027
Chisq
0.13146
NA
R
0.98545
NA
4.5
4.0
160
170
180
190
200
210
220
230
Power (W)
FIGURE 3.4: a) Transition flow and power values for depositions from one Al target.
b) Transition flow vs. corresponding power level (from graph a).
Argon flow: 80sccm. Current: 0.80 A. Each deposition lasts approximately two hours.
The controlled variable is the transition oxygen flow, QT.
41
This speculation is supported by the drift in the initial voltage values Vinit, as the
target is used repeatedly for reactive deposition of aluminum oxide films. In
effect, we observed a dependence which can be described as:
Vinit = f (n) = Vinit (n)
where n is the number of depositions performed by a single target. To avoid any
effects caused by the exposure of the target to the atmosphere, we compared the
discharge voltage values before (i.e.: Vinit) and after each deposition sequence.
Vinit was 2-10 Volts higher than the voltage measured in a pure argon plasma
immediately after the process. Also, in the following reactive deposition
sequence, Vinit(n+1) was lower than Vinit(n). Consequently, the new transition
voltage, empirically defined as Vinit-30, decreased too.
The most useful conclusion from Figure 3.4 is the linear relation between the
average power of the discharge and the oxygen flow when the target is at the
transition state. This relation has been used to predict the transition flow for
reactive depositions with different aluminum targets with just the knowledge of
the discharge voltage Vinit in a pure argon atmosphere(Figure 3.5). In particular,
the desired level of the average power during the reactive deposition is
calculated as:
S
avg
= 0.8 × ( V init – 30 )
(3.1)
and then the predicted transition flow (based on the linear regression analysis on
the data in Figure 3.4.b) is given by:
Q T = ( – 2.26 ± 0.52 ) + ( 0.0383 ± 0.0027 ) × S
avg
(3.2)
42
Of course, this is just a guide for the correct range of the oxygen flow. The actual
value, which will keep the voltage and power at the desired levels, will have to
be dynamically re-adjusted during the deposition.
Transition O2 flow (sccm)
6.5
6.0
5.5
5.0
Q = m1 + m2 * S
T
4.5
4.0
160
170
180
Value
Error
m1
-2.2602
0.52098
m2
0.038283
0.0026958
190
200
210
220
230
240
Power (W)
FIGURE 3.5: Actual (markers) and predicted (line) transition flows for 4 depositions.
Predicted transition flow values are derived from Eq.(3.2).
The variance in the deposition rates is a direct result of the transition point drift,
which depends on the history of target. The power of the discharge is directly
related – above a threshold and below a saturation value – to the deposition rate.
Another change that occurs with repeated and prolonged usage is in the shape of
the target. As the target is sputtered, a circular erosion track forms on its surface,
along the field lines created by the permanent magnets in the magnetron gun.
This track has grown as deep as 2mm in our targets, which have a total thickness
of 3mm. It may be argued that the increased effective surface of the target causes
a decrease in the current density and the voltage required to supply the constant
current of 0.80 A.
It has been suggested that prolonged pre-sputtering of the “poisoned” target in a
pure argon plasma with a dc voltage might ultimately “clean” the target and
restore Vinit to its value for a fresh target (~310V, at 0.80Pa, 0.80A, 181kHz).
43
However, it was not possible to achieve such results after 30 minutes of
dc-sputtering in pure argon. In fact, when the target was driven with the
standard pulsed dc signal, the discharge voltage remained the same as before
pre-sputtering.
The fluctuation of the oxygen flow and the discharge power (or voltage) during a
single deposition sequence, as depicted by the error bars in Figure 3.4 is
attributed to two causes. First, there is a drift in the measured flow level with
time. The measured flow value will be decreased by up to 0.4sccm in the course
of a 2 hour long deposition, even though the set point in the mass flow controller
has remained at the same value. This drift can cause the subsequent increase of
the discharge voltage towards values higher than the transition point. Then, the
operator may adjust – generally, increase – the oxygen flow so as to maintain the
voltage at the transition level. Furthermore, due to the inherent instability of the
process, a momentary perturbation in the flow may send the voltage irreversibly
up or down the curves shown in Figure 3.3. At that point the operator has to
adjust the flow in order to return the voltage to the transition value. In the case
where the voltage has reached its lower limit, the deposition has to be
interrupted and the target must be sputtered in a pure argon plasma so that the
discharge voltage reaches, or at least approaches, Vinit.
From the discussion above it may have become obvious that it is difficult to
predict the transition oxygen flow for each deposition. The linear fit from
Figure 3.4 can be used to estimate the transition flow QT with relative accuracy in
only half of the cases that were used to derive it. It can certainly be employed to
obtain a rough estimate but the fine-tuning of the process must be performed
either by the operator or by an “expert” system which would actively control the
oxygen flow using the power values as a feedback.
44
Discussion of the relation between discharge power and deposition rate
As has been discussed earlier, the discharge power for each deposition sequence
is determined by the voltage Vinit of the target in a pure argon plasma. This initial
voltage is decreased as the target is used repeatedly. Consequently, the oxygen
flow that is necessary to bring the target to its transition state, empirically
defined as the state where V = VT = Vinit-30, decreases also.
It is expected that as the power of the discharge is decreased from one deposition
to the next, the deposition rate will also suffer a similar decline. In order to
substantiate this the rate at which the oxygen was consumed in each deposition
was estimated. This estimate is based on the schematic in Figure 3.6.b, where the
change of total pressure with oxygen flow is depicted with and without plasma
discharge. A model with a finite transition range of flows is used in order to
better approximate the experimental reality.
If an average transition flow of oxygen QT is assigned for the total duration of a
deposition sequence then, in principle, it is possible to estimate the partial
pressure of oxygen corresponding to that flow, if the former is calibrated against
the latter. This calibration was done by measuring the partial pressure of oxygen
for various flow values in the 4-6.5sccm range. The particular interval was
selected because in almost all depositions we selected a transition flow within
that range. During all measurements a constant argon flow of 80sccm was kept
and the main pump was throttled to maintain an argon partial pressure of
0.80Pa. Unfortunately, the stability of the latter cannot be guaranteed during the
measurements since the pumping speed drifts with time – generally decreases –
even when the flow remains constant. Moreover, the pumping speed is not
constant at the pressure levels resulting from mass flows of tens of sccm.
The exact values of our measurements and the result of a linear regression
analysis are presented in Figure 3.6.a. The continuous pressure curve in
45
Figure 3.6.b is essentially the linear fit curve derived from the pressure
measurements.
(a)
(b)
0.9
0.1
0.06
Value
Error
P00 -0.0017
0.0021
a
0.0123
0.0004
R
0.9850
NA
0.88
0.86
0.04
0.84
0.02
0.82
0
-0.02
1
2
0.8
P0
P00
0
No discharge
Discharge with increasing flow
+a*Q
3
4
5
6
7
8
Total pressure (Pa)
Partial pressure (Pa)
0.08
00
Total pressure
P=P
Metallic mode
Mode A
Pn
Insulator mode
Mode B
PT
T
PAr
0.78
Oxygen flow (sccm)
Q
T
Oxygen flow
FIGURE 3.6: a) Calculation of partial pressure–oxygen flow relation.
b) Total pressure vs. increasing oxygen flow with and without discharge.
Graph (a): Argon flow: 80sccm. Argon partial pressure: 0.80Pa. No discharge.
Graph (b): A realistic case with a finite transition range of flows is depicted. Constant
argon flow and partial pressure. Constant pumping speed.
From Figure 3.6.b it is evident that at a flow equal to QT only the surplus oxygen
contributes to the total pressure and its partial pressure is just PT-PAr, where PT is
measured directly with a capacitance manometer, and PAr is measured in a pure
argon plasma with 80sccm of argon, just after the deposition sequence, at the
same throttling of the pump. Naturally, the consumed oxygen is responsible for
the missing pressure which, if added to PT, would bring the total pressure to Pn.
Then, the fraction of the supplied oxygen that is consumed is given by:
( P n – P Ar ) – ( P T – P Ar )
f = --------------------------------------------------------------P n – P Ar
(3.3)
46
However, Pn cannot be measured at the same conditions as PT. It can only be
calculated from the linear fit from Figure 3.6.a:
P n = P 0 – 0.0017 + 0.0123 × Q T
(3.4)
where P0 was adjusted to be 0.8Pa and represents a constant pumping speed. It
has to be noted that P00 (see Figure 3.6.a) should be equal to zero
(0.0000±0.0025Pa) if the pumping speed was constant in the pressure range
encountered during calibration. Its calculated value (-0.0017±0.0021Pa) comes
close enough, yet its slightly negative value is evidence that the pumping speed
at 0.84-0.88Pa is lower than that at 0.80Pa, and therefore the pressure-flow curve
at 0.84-0.88Pa is steeper.
PAr and P0 are measured in different conditions – with and without discharge,
respectively – and therefore there is a leap of faith in accepting the partial
pressure of oxygen without discharge as the difference Pn - PAr. To the extent that
P0 can be approximated with PAr and inasmuch the pumping speed is constant
during calibration and deposition, we can use Eq.(3.3) with relative confidence,
keeping always in mind the uncertainties in pressure readings and mainly the
drift in the apparent pumping speed.
As stated before, of interest is the consumption rate of oxygen, which is just the
product of the supplied rate with the fraction that is consumed:
Pn – PT
C r = Q T × f = Q T × ----------------------P n – P Ar
(3.5)
The variation of Cr for two targets as a function of the average discharge power
utilized in each deposition is presented in Figure 3.7. It is observed that during
the useful life of a particular target, which is always sputtered at the transition
voltage VT = Vinit-30, the estimated consumption rate of oxygen suffers a gradual
47
decrease of 45%, even though the supplied transition flow of oxygen is reduced
by less than 35% (Figure 3.4.b). This difference implies that the consumed
fraction of the supplied oxygen decreases throughout the life of a target, even
when the discharge voltage is adjusted to the empirically determined transition
value.
Another significant observation is that the decrease in the estimated
consumption rate of oxygen explains the measured reduction in the deposition
rate of the aluminum oxide. The latter falls from 0.7nm/s for a fresh target to
0.4nm/s towards the end of the target life. This agreement substantiates the
preceding analysis since a high consumption rate of oxygen, in the transition
state, will result to a high production rate of aluminum oxide and consequently
to a high deposition rate.
Average power (% of maximum)
0
60
80
100
T
1.2 10
(100 = 1.15 x 10
18
init
VT=V init-30
1 1018
80
8 10
18
mc/s)
40
V =V -15...85
100
Oxygen consumption rate
20
17
60
6 1017
40
20
0
0
50
100
150
200
4 10
17
2 10
17
0
250
Average power (W)
FIGURE 3.7: Oxygen consumption rate vs. discharge power.
Closed circles: depositions from one target. Open circles correspond to depositions
from different target, carried out in different conditions (see legend).
Finally, the aberrant value in Figure 3.7 which corresponds to a power of 150W
represents a deposition sequence where the target was fully oxidized (due to a
high oxygen flow rate). In this case the fraction of the oxygen that is consumed is
48
so low that reduces dramatically the consumption rate despite the high supplied
flow.
3.2.2 Deposition sequence
Ordinarily, it is necessary to deposit an insulating layer over a part of the
substrate only, the goal being to minimize the area occupied by an insulator at
the interface between the protective layers and the substrate. Thus, an increased
mechanical integrity is achieved as the metallic protective layer is being grown
on top of a metallic body. The minimization of the obstacle to the heat diffusion
and conduction that is imposed by the insulating film is also achieved. Currently,
a simple shadow masking technique is employed. In particular, a piece of thin
304L stainless steel sheet, having the overall dimensions of the substrate, is
mounted on top of the substrate on the substrate holder. The piece of sheet has a
machined slot so as to leave part of the substrate exposed to the impinging
molecules. In Figure 3.8 such a piece, used for 50x50mm substrates, is depicted.
FIGURE 3.8: Shadow mask for shaping sputter-deposited Al2O3 layers.
The deposition of the bottom insulation layer is preceded by the deposition of an
adhesion-promoting titanium film. The latter takes place immediately after the
completion of the sputter-etching process which has been described in “Sputteretching” on page 31. It is important to note that sputter-etching, titanium and
aluminum oxide deposition, all occur in the same vacuum in order to minimize
contamination of the surfaces. The parameters used for the deposition of the two
films are described in Table 3.8.
49
TABLE 3.8: Process parameters for reactive deposition of aluminum oxide.
Target material
Titanium
Aluminum
Base pressure
<1E-4Pa
N/A
Argon flow
80±2sccm
80±2sccm
Argon partial pressure
0.80±0.01Pa
0.8000±0.0025Pa
Oxygen flowa
N/A
4-6.5sccm
Oxygen partial pressure
N/A
0.3-0.6Pa
Power or Current
(controlled)
150W dc
0.80Ab
Pulse frequency/width
N/A
181 kHz/1056nsc
Discharge voltage
310-320V
210-280V
Deposition time
20-30s
5000-7000s
Film thickness
10-20nm
3-5µm
a. Following the “Discussion of the history dependence of the transition point” on
page 39, it is evident that, for each particular deposition sequence, the oxygen flow has
to be adjusted in order to achieve a discharge voltage close to VT which is defined as
the voltage 30 Volts below the discharge voltage in a pure argon atmosphere (Vinit).
b. Value set at 80% of maximum.
c. Lower frequency and pulse width values lead to long-term plasma instabilities, while
higher values reduce the apparent deposition rate.
The substrate is mounted on the holder so as to be directly across the target
during the deposition. However, due to the small target size (diameter of 50mm)
it is impossible to get very uniform deposits with stationary substrates. Usually,
the thickness profile of thick films (2-3µm) has a slope of 200-300nm across a
radial distance of 1cm.
The deposition of the titanium film is quite straightforward. There is a 300s presputtering period in order to remove the oxide and adsorbed water molecules
from the target surface and to achieve a stabilized discharge voltage. During this
stage, the substrate is shielded with a shutter. The removal of the shutter marks
the start of deposition.
50
Before every reactive deposition the aluminum target is also pre-sputtered for
300-600s in a pure argon atmosphere in order to remove surface contamination
(region 1 in Figure 3.9). During this step, the shutter is placed between the
substrate holder and the target. At the final stage (60-100s) of pre-sputtering, the
substrate holder and the shutter are moved away from the target in order to
simulate better the conditions during actual deposition. The reason is the
influence of the target-shutter distance (3cm) to the discharge voltage. By
removing the shutter – and, of course, the substrate holder behind it – the
distance from the target to the nearest obstacle (15cm) is significantly increased
and therefore more ionization events are allowed to occur, which in turn decrease
the discharge voltage necessary to sustain the pre-set current level of 0.80A. In
this way a more reliable value for the initial voltage Vinit can be obtained and will
be used to define the transition voltage VT. During this stage, the turbomolecular
pump is throttled so as to achieve a pressure of 0.8Pa.
In order to enhance the “atomic peening” effect [cf. “Stress state” on page 14], a
number of deposition experiments were carried out at partial argon pressures of
0.4-0.5Pa. However, it was impossible to sustain the discharge over a few
minutes without extensive arcing. Therefore, the partial pressure of argon was
kept at 0.8Pa for all subsequent experiments.
The next step is to introduce oxygen at a level that will cause the discharge
voltage to drop by 30 Volts (region 2 in Figure 3.9). An estimate for the necessary
oxygen flow can be made using the linear fit from Figure 3.4.b. which relates the
discharge power to the average oxygen flow. However, it is necessary to finetune the flow in order to get a stable voltage before the return of the substrate
holder at the appropriate position for deposition (facing the target). This replacement of the substrate holder marks the start of the deposition sequence
(region 3 in Figure 3.9).
51
During the entire deposition the discharge voltage should be monitored and any
drifts from the desired value (VT) should be compensated by adjusting the flow
(circled flow markers in Figure 3.9). In particular, when the voltage drifts to
higher values, the oxygen flow should be increased in order to bring the target to
a more “poisoned” state and counter the voltage hike. The reverse action is taken
in the case of a voltage decrease.
(1)
260
(2)
(3)
5.2
Flow adjustments
O2 Flow (sccm)
250
5.0
Voltage (V)
Flow drifts
4.8
230
220
4.6
210
4.4
Voltage drifts
200
O Flow (sccm)
2
240
7300
7200
6900
6740
6640
6400
2720
2180
2000
1870
930
1780
750
650
560
510
460
420
280
90
190
190
Voltage (V)
4.2
Time (s)
FIGURE 3.9: Discharge voltage and oxygen flow variations during reactive deposition.
Region 1: pre-sputtering; region 2: initial oxygen injection; region 3: reactive deposition
of aluminum oxide.
Notice the drifts in the discharge voltage and the flow adjustments employed to
counteract them.
It might happen that the voltage has drifted to its low limit before the operator
takes notice of the change. In such a case an incremental flow decrease will force
the discharge voltage to trace the lower (dashed) curve of the hysteresis loops in
Figure 3.3, and the transition point will be reached at a lower flow. This is what
happened in the case of Figure 3.9: initially QT was 5.1sccm. After the voltage
had drifted to its low limit we decreased incrementally the flow until we raised
52
the discharge voltage to VT. However, the new QT was only 4.5sccm. In order to
perform the deposition at the initial conditions we would have to shield the
substrate with the shutter, interrupt the oxygen flow, wait until the discharge
voltage returns to within 2-3 Volts from Vinit, and then re-inject oxygen at the
original flow level and resume deposition.
3.3 Sensor layer
The materials used for the sensor circuits were selected with respect to their
thermo-electric properties. In particular, strain gages are sputter-deposited from
a constantan target because this alloy has a very low temperature coefficient of
resistivity (30ppm @ 25-105°C). Also, the standard thermocouple alloys or metals
(alumel, chromel, constantan, copper) were used to build temperature sensors by
sputtering the respective targets. The obvious reason for this selection is the ease
in calibrating the deposited sensors using standard thermocouple reference
tables.
3.3.1 Photolithography
The sensors are in essence electrical circuits and therefore they have to be shaped
out of the blanket metallic film that is deposited via sputtering. The shaping is
achieved with micromachining which combines photolithography (steps (a)-(d)
in Figure 3.10) and a removal sequence known as lift-off (step (f) in Figure 3.10).
The main goal of this work has been to establish the feasibility of embedding
sensors in metallic structures and therefore there was not put much of an
emphasis on the geometry of the circuits. The designs (positive masks) for the
strain gages are included in the Appendices (“Positive masks of strain sensors”
on page 124). The line width varied from 420 to 50µm and the size of the longest
line varied from 30 to 8mm. The line width for the thermocouple was generally
176µm and the length of the sensor itself was 30mm.
53
a) Photoresist is deposited
on substrate
b) Patterned mask is placed
on substrate and exposed
c) Mask is removed
d) Exposed photoresist is developed,
leaving a patterned window
e) Metal thin film is deposited
over the substrate
f) Photoresist layer is removed by acetone,
carrying away overlying metal film (Lift-off)
FIGURE 3.10: Schematic of the photolithography steps preceding a lift-off process.
The masks were designed with the aid of a commercial drawing application,
usually Adobe Illustrator™, and after their transfer to a Postcript™ file the
negative of the image was printed on a transparency by a high-resolution laser
printer.
Then, the mask pattern was transferred on the substrate surface in the following
manner: a few drops of positive photoresist (Shipley 1813) were dropped on the
substrate which was spinning at 2000-3000rpm for 30-40s. The rapid spin
resulted in a thin (1.5-2.0µm) layer of photoresist. The substrate was
subsequently baked in an isothermal furnace at 90°C for 20 min. After baking,
the part of the mask with the desired sensor was positioned manually on top of
the substrate and was held in place by a glass plate. The glass plate was pressed
54
against the substrate by a simple spring-loaded assembly. The assembly was then
placed under an UV light source and was exposed for 40 seconds. Exposure of a
positive photoresist to UV radiation causes a change in the chemical structure of
the photoactive agent of the resist. This change makes the exposed volume
soluble to developer solutions.
After exposure the substrate was removed from the assembly and was immersed
for 15-20 seconds in a bath containing concentrated developer from Shipley and
de-ionized water in a 1:1 ratio. The developing time is enough to remove the
photoresist from the exposed regions which correspond to the sensor pattern.
Developing is stopped by rinsing in de-ionized water.
Stylus profilometry of the developed region near the sidewalls has shown that
the thickness of the photoresist is reduced from its maximum value (1.5-2.0µm)
to zero in a distance roughly equal to 8µm. This corresponds to a rather shallow
slope, suboptimal for the lift-off process which requires ideally negative slopes or
at least vertical ones. Nevertheless, the lift-off results for the particular line
widths have been quite successful.
3.3.2 Deposition
The substrate is mounted on the substrate holder and transferred to the
sputtering chamber. After the pump-down the deposition sequence begins with
the growth of a thin (10-20nm) adhesion layer out of titanium. Then the substrate
is transported over the magnetron gun with the sensor alloy target and the
deposition of the sensor layer commences. The parameters for the deposition of
each particular film are summarized in Table 3.9.
The sputter-etching of the substrate is generally not recommended in order to
prevent damage on the dielectric layer and the extreme hard-baking of the
photoresist. The substrate is bombarded with high energy (~300eV) ions and
electrons (since a pulsed-dc signal is used to drive the substrate holder) during
55
the sputter-etching process, and its temperature is raised to ~100°C, thereby any
S1813 photoresist layer would be unintentionally hard-baked and the lift-off
procedure would be considerably hampered.
TABLE 3.9: Deposition parameters for the sensor alloys and the adhesion interlayer.
Target material
Titanium
Constantan
Alumel
Chromel
Base pressure
<1E-4Pa
<1E-4Pa
N/A
N/A
Argon flow
80±2sccm
80±2sccm
80±2sccm
80±2sccm
Argon partial
pressure
0.80±0.01Pa
0.80±0.01Pa
0.80±0.01Pa
0.80±0.01Pa
Power
(controlled)
150W dc
150W dc
150W dc
150W dc
Discharge
voltage
310-320V
390V
410-420V
408V
3.3.3 Lift-off
As soon as the sensor layer is deposited, the substrate is removed from the
chamber and immersed in acetone. It is left to soak for 4-5 hours, a period which
is usually adequate for the photoresist be dissolved. As the photoresist is
removed, the metallic films that have been deposited on top of it lose their base
and can be washed by the acetone. Only the region of the films that was
deposited on top of the dielectric, through the window in the patterned
photoresist, remains bonded to the substrate and forms the sensor.
Ideally, in the case of vertical or negative photoresist sidewalls, the film
deposited on top of the photoresist would be fully detached from the part which
is bonded to the substrate. However, in the case of positive slopes, the film
remains attached to the bonded region along its edges. The unwanted film is
removed by “tearing” these edges with the aid of pressurized dry nitrogen. The
result is a relatively jagged edge on the sensor film as shown in Figure 3.11. This
undesired effect can be eliminated by using either negative photoresist (where
56
FIGURE 3.11: Jagged edge of deposited constantan film after lift-off.
the exposed part stays, while the unexposed is dissolved), or a combination of
chemical treatment and exposure that reverses the slope of the positive
photoresist walls.
3.4 Top Oxide
The deposition sequence for the passivation layer that partly covers the sensor is
almost identical to that for the base insulation layer. The only differences are
relevant to masking and sputter-etching.
Since it is necessary to leave parts of the sensor exposed for splicing purposes
(splicing is the connection of wiring to the sensor to enable signal acquisition and
processing) it is apparent that the mask in Figure 3.8 cannot be used as is.
Usually, a piece of aluminum foil is used to cover the connection pads of the
sensor during deposition. The rest of the substrate is covered with the shadow
mask used in the deposition sequence of the first oxide layer.
57
Sputter-etching is critical in achieving a pinhole-free oxide layer by removing the
top few monolayers of the substrate where dust particles or organic molecules
may be attached. However, in this case the sensor film itself would be thinned by
the sputter-etching action and its resistance would change. This is not an issue
for the thermocouples where the output is only dependent on the difference of
the thermoelectric coefficients of the two materials. Nevertheless, the
generalization of a strain sensor calibration for a batch of similar devices would
only be valid if the sputter-etching process altered the thickness and the
resistance of the sensor in a well-defined and predictable manner.
3.5 Protective Layers
The purpose of these layers is to protect the thin film structure (insulation and
sensor) from the high-temperature embedding process. In particular, the
protective layers are necessary to reduce the temperature experienced by the thin
films as an intense and localized heat flux is imparted by the laser during the
formation of the embedding layer. The results of an attempt to form such a layer
directly over the oxide films can be seen in Figure 3.12.
Initially, it was thought that a relatively thick (2-3mm) layer of copper (a metal
with a thermal conductivity second only to silver) could be used to rapidly
transfer the heat generated during the embedding step to the metal substrate. A
necessary provision for this layer would be to directly contact the substrate over
a large area, compared to the surface occupied by the dielectric films that were
sandwiched between the substrate and the protective layer (Figure 3.14.b). The
thick copper layer could be grown by electrodeposition on top of a “seed” copper
film which would be sputter-deposited on the substrate, partially covering the
top dielectric layer.
Embedding experiments showed that the copper layer could not function as the
sole protective coating for a number of reasons. First, its high thermal
58
Laser deposited invar
Al2O3
2 mm
304L stainless steel
substrate
FIGURE 3.12: Cross-section showing an Al2O3 film damaged during laser deposition.
Notice the discontinuity in the thin film and the remelted invar/304L interface between
the oxide fragments.
conductivity would not allow partial remelting near its surface, thereby causing
the formation of void clusters at the interface with the embedding layer as shown
in Figure 3.13.a. Such a condition would severely compromise the mechanical
integrity and the fatigue resistance of the structure. In addition, even in the cases
where partial remelting was achieved, the copper and the invar layer would
retain their distinct phases, as can be seen in Figure 3.13.b, thereby falling short
of a strong metallurgical bond.
So, it became evident that a second protective layer could be essential for the
successful embedding. Bearing in mind that electrodeposition would be the
process of choice, we focused on a small number of elements and selected nickel
for its lower thermal conductivity and the fact that it is one of the two
constituents of the embedding alloy (invar). In the new configuration, the nickel
layer would cover the copper layer and it would extend over the edges of the
latter in three sides, as shown in Figure 3.14.c.
59
10 µm
50 µm
Electroplated Cu
(a)
Electroplated Cu
(b)
FIGURE 3.13: a) Clustered voids at the invar/copper interface.
b) Distinct phases across the partially remelted invar/copper interface.
(a)
(b)
(c)
Top dielectric film
Sensor film (contact pads)
FIGURE 3.14: a) Schematic of substrate with sensor enclosed between two dielectric
films.
b) Schematic showing the copper layer footprint.
c) Schematic showing the nickel layer footprint.
3.5.1 Copper layer
Since an electroplated layer can only be grown on a conductive substrate, it was
necessary to cover the top dielectric film with a metallic layer which would serve
as a “seed” for the electrodeposited copper. Sputter-deposition is a relatively
simple process that can deposit metallic films on a great variety of substrates and
was selected to grow a thin film of copper on the substrate. A titanium interlayer
was deposited in order to enhance the adhesion of copper to both the aluminum
oxide and stainless steel areas. The deposition of the two films was sequential
60
and was carried out in the same vacuum. The process parameters for the sputterdeposition are described in detail in Table 3.10.
TABLE 3.10: Deposition parameters for the “seed” copper film and the adhesion
interlayer.
Target material
Titanium
Copper
Base pressure
<1E-4Pa
N/A
Argon flow
80±2sccm
80±2sccm
Argon partial pressure
0.80±0.01Pa
0.80±0.01Pa
Power
(controlled)
150W dc
150W dc
Discharge voltage
310-320V
400-410V
Film thickness
10-20nm
1-1.2µm
To simplify the process, the shaping of the “seed” film was achieved by simple
shadow masking. Obviously, a more exact, photolithographic procedure can be
used if necessary.
After the growth of the “seed” layer, the area of the substrate that would not be
covered by electroplated copper was masked with a relatively thick (3-4µm)
layer of photoresist. The back side of the substrate was fully covered with kapton
(polyimide) tape or a thick resin, except for a small area that would be used for
electrical contact. Finally, a die-like wall structure was placed around the area
that would be electroplated. Generally, specially shaped kapton tape or teflon
was used in order to create a 4-6mm tall wall to contain the deposit up to its total
thickness of 1-2mm.
Special care was taken during the handling of the substrate so that the sputterdeposited “seed” film of copper would not come in contact with anything but deionized water and that the whole preparation sequence, from the time the
substrate was exposed to air until the time of its immersion in the electroplating
61
bath, did not take more than 1 hour. The reason for this precaution was to
minimize the oxidation of the sputter-deposited copper surface, as this would
affect in an adverse manner the adhesion of the electroplated layer to the “seed”
film.
In the few cases of mishandling that caused the formation of a surface oxide on
the copper film, the substrate was immersed for a few seconds in a solution
containing 10% v/v HCl. The acid would remove the oxide without attacking the
metallic copper. However, such processing involved the risk of acid reaching the
buried top oxide film through pinholes in the copper film. Since the aluminum
oxide was deposited at a very low homologous temperature (Tsub /Tm<0.2), it was
purely amorphous [cf. “Microstructure” on page 79] and therefore susceptible to
etching by the acid. Therefore by allowing the dielectric film to be attacked, there
was a risk of exposing the encapsulated sensor lines (through tiny pinholes) to
the electroplating bath and causing them to short-circuit to the protective layer.
Such an event would severely harm the functionality of the sensor. Since it was
observed that in samples treated with HCl for the removal of possible surface
oxides the insulation of the sensors was compromised after only minutes of
electroplating, this procedure was actively avoided.
As soon as the substrate was ready, it was connected to the live cathode of the
power supply in the electroplating system and then it was lowered into the bath
which was generally agitated by air flow.
Copper can be electrodeposited either in a cyanide or an acid bath. For reasons of
operation safety the more benign acid method was chosen and its chemical
composition is described in Table 3.11.
62
TABLE 3.11: Chemical composition of copper electroplating solution.
Copper sulfate pentahydrate
220-240g/l
Sulfuric acid
55-65g/l
Chlorine ions
30ppm
Aluminum potassium sulfate dodecahydrate
2g/l
The exact parameters used during the electrodeposition process are presented in
detail in Table 3.12.
TABLE 3.12: Electrodeposition parameters for the copper protective layer.
Initial current density
21mA/cm2
Time
1hr
Final current density
31.5mA/cm2
Time
30hr
Total thickness
~1mm
Temperature
27°C
At the end of the deposition the substrate was removed from the bath, rinsed
with distilled water and blown dry after the removal of the die-like wall
structure.
3.5.2 Nickel layer
Initially, the exposed polished stainless steel surface was mechanically
roughened via grinding in order to improve the mechanical interlocking factor
with the nickel layer. Subsequently, part of the surface was covered with a die
structure surrounding the area to be electroplated, defined by the outer dashed
line in Figure 3.14.c, in order to contain the thick nickel deposit.
The deposition of the nickel layer was performed in two steps also. This time the
“seed” layer was grown in a nickel “strike” bath, on top of the copper layer and
part of the exposed stainless steel substrate. Prior to the “seed” layer deposition
the substrate was subjected to electropolishing by the application of a reverse
voltage (substrate connected to the anode) for 2 minutes. Then the voltage was
63
reinstated to its forward direction and the deposition commenced. The
composition of the “strike” bath and the process parameters are presented in
Table 3.13 and Table 3.14 respectively.
TABLE 3.13: Chemical composition of nickel “strike” electroplating solution.
Nickel chloride hexahydrate
60g/l
Hydrochloric acid
125ml/l
TABLE 3.14: Electrodeposition parameters for the “seed” nickel layer.
Reverse current density
21mA/cm2
Time
2min
Forward current density
21mA/cm2
Time
2min
Total thickness
~1µm
Temperature
27°C
The seed layer could also be deposited via sputter-deposition. However, that
would involve an additional pump-down sequence and increased exposure to
the air during the transfer between the chamber and the electroplating bath.
Using a “strike” solution instead, the process time was minimized and the
substrate is transferred immediately to the next bath while it is still wet.
Following the “strike” deposition, the substrate was transferred immediately to
the nickel sulfamate bath for the growth of the nickel layer. That bath was acidic
in nature and its chemical composition is listed in Table 3.15, while the
deposition process parameters are presented in Table 3.16.
TABLE 3.15: Chemical composition of nickel sulfamate electroplating solution.
Nickel sulfamate
265ml/l
Boric acid
38.6g/l
Barrett additive-A
3.8g/l
Barrett SNAP A/M
0.3% by vol.
64
TABLE 3.16: Electrodeposition parameters for the nickel protective layer.
Initial current density
21mA/cm2
Time
2-3hr
Final current density
42mA/cm2
Time
40-50hr
Total thickness
1-1.5mm
Temperature
48°C
At the end of the deposition sequence, the substrate was retracted from the
solution and all the protective insulating layers (photoresist, resin, kapton tape)
were removed. Finally, the substrate was rinsed in distilled water and blown dry.
3.6 Embedding Layers
The final embedding of the sensor is accomplished with a high-temperature
process. In the Rapid Prototyping Laboratory, there are two options available for
producing fully dense metal layers: a) plasma microcasting, and b) laser
deposition. The former uses a plasma arc to superheat the tip of a metallic wire in
order to produce a droplet that subsequently falls onto the substrate, where it
solidifies upon impact. By moving the wire relative to the substrate (or vice
versa) these solidified droplets can form a dense layer of material. The latter
method (cf. “Laser Assisted Deposition” on page 23) uses a high power (2.4kW)
infrared Nd:YAG laser beam to create a molten pool on a metal powder bed
which is pre- (or concurrently) deposited on the substrate.
The beam is created in the optical cavity and transferred via an optical fiber
whose end is attached to a computer-controlled robotic arm. The material in the
spot of the laser beam is heated beyond its melting point and forms a molten
pool. By moving the beam relative to the substrate the molten powder cools
rapidly and is fused to the underlying material, thereby producing a bead of
dense material.
65
The laser deposition technique was chosen over plasma microcasting due to
superior positioning control of the deposit. Indeed, plasma microcasting suffered
from poor control in the landing position of the droplet and its deposit was
rather shapeless due to the splattering of the droplet upon its collision with the
substrate.
3.6.1 Thermal stress minimization
Initially, all embedding experiments were conducted with the substrate lying on
top of a large aluminum mass, acting as a heat sink. The results were rather
discouraging as a large percentage of the samples suffered from delamination
between the embedding layers and the underlying structure. The delamination
was caused by the large thermal stresses developed during the process. Since the
thermal stresses originate from thermal strains, the minimization of the latter will
lead to the minimization of the former (for the same materials). In layered multimaterial structures the thermal strains are caused by differences in the thermal
expansion coefficients (α) as well as temperature differences among the
constituting layers. The resulting thermal stresses will be accentuated by the
elastic moduli of those layers. For a given set of substrate materials the only
degrees of freedom lie in the selection of the last (topmost) layer and the
temperature range of the processing.
Therefore, the selected material for the embedding layer should have the smallest
possible thermal expansion coefficient, elastic modulus and melting temperature
in order to ensure minimal thermal stress generation. However, since the goal of
the process is to produce a structure that can withstand high temperatures, it is
obvious that the melting point and, consequently, the elastic modulus, must have
high values. This reasoning leaves α as the sole variable that can be minimized
by material selection. The only high-meting point metal that offers a distinctly
small α is invar (Fe0.64-Ni0.36).
66
With respect to the temperature gradient experienced by the structure the lower
bound of the temperature at the topmost layer is the melting point of the
embedding layer. It was assumed that the laser beam would raster the powder
layer at an optimal speed, sufficient to produce a melting pool as deep as the
powder bed itself. Lower speeds would unnecessarily increase the heat flux to
the structure, whereas higher speeds would leave unfused powder behind the
laser path. Therefore, the only way to decrease the temperature gradient would
be to increase the initial temperature of the substrate. Such a move would also
decrease, by a little, the necessary heat input to achieve fusion of the powder
layer.
So, the 50mm square stainless steel substrates were placed on top of a copper
slab (20 x 20 x 1cm), which in turn was sitting on top of a temperature-controlled
hot plate. The latter was set at 450°C and the copper slab would reach thermal
equilibrium at ~350°C. In order to provide a consistent thermal interface between
the stainless steel substrate and the copper plate, irrespective of the surface finish
and the powder particles that might be attached to the contacting surfaces, a thin
layer of silicone paste was applied to the bottom the stainless substrate. The
equilibrium temperature of the substrate prior to deposition was 300-350°C.
3.6.2 Fusion of the powder layer
The powder was placed was placed either automatically via a feed tube which
was scanned above the substrate by a robotic arm or manually by the operator.
The thickness of the powder layer was adjusted to 3±0.5mm. A “supporting”
mass of powder was also arranged around the substrate in order to allow
constant thickness of the powder layer all over the substrate surface.
The attributes of the laser path, i.e.:coördinates, scanning speed, intervals
between passes, were gradually determined by making educated guesses based
on observation of attempts which succeeded only partially or failed
catastrophically. Another decision which was influenced by the observation of
67
such attempts concerned the addition of an extra protective layer. The
progressive attempts to arrive at final and successful embedding procedure are
depicted in Figure 3.15 and Figure 3.17.
During the initial attempts the laser beam was scanned in a zig-zag fashion over
the powder bed, with a constant linear speed of 20mm/s and without a stop
between successive passes (which would be traced in opposite directions). The
power was held constant at ~1.8kW and the diameter of the beam spot was
2.5mm, yielding a power flux of ~360MW/m2. The result of this configuration
when used on a non-pre-heated sample with just one protective copper layer was
inadequate bonding of the invar at the beginning of the path and through-thethickness remelting of the copper near the end of the path.
2
1
10
(a)
1
2
10
(b)
(c)
FIGURE 3.15: Laser path during progressive attempts to improve embedding.
a) Original, continuous path; circle signals the end of the path;
b) 20 second intervals introduced between passes; numbers signify pass order;
c) Last four passes carried out at higher speeds.
This clearly indicated a problem with the amount and rate of heat that was
injected to the sample. A first response was to introduce 20 second intervals at
the end of each path, allowing for the substrate to cool down. When this strategy
was used in conjunction with pre-heating the substrate to a temperature of
~350°C, the sample was cooling during those intervals to within a few degrees
from its equilibrium temperature. In addition to switching off the laser power,
another measure was to reduce the total amount of heat injected to the sample by
68
accelerating the scanning speed towards the end of the path. So, during the last 4
passes of the path the laser was moved at a linear speed of 20-25mm/s, instead
of 15mm/s.
All these improvements did away with the remelting of the copper layer through
its thickness near the end of the path. However, two other issues remained
unresolved; namely, the inadequate bonding of the fused invar layer to the
copper (Figure 3.13.a) and the delamination of the copper layer from the
substrate at the edge near the end of the path (dotted edge in Figure 3.15.c). The
former issue was attacked successfully only with the introduction of a nickel
protective layer. The improvement in the quality of the interface can be seen in
Figure 3.16 below.
50 µm
Electroplated nickel
FIGURE 3.16: Invar-nickel interface from a cross-section of an embedded structure.
The localized delamination of the copper layer was addressed by a modification
in the path of the laser beam and the introduction of substrate pre-heating. The
latter increased the temperature of the stainless steel substrate, thereby
decreasing the magnitude of the ∆α∆Τ product for the copper-stainless bilayer.
The laser path was altered so as to fuse the powder initially over the two edges of
the protective layers (passes #1 and #2 in Figure 3.17) and subsequently scan the
69
rest of the powder bed in a zig-zag fashion.The actual LabView code that
produces the path in Figure 3.17 is included in page 125, along with comments
for the various commands.
4
2
3
1
11
FIGURE 3.17: Final laser path configuration.
The numbers in the arrowheads signify the pass order. The dashed lines stand for
higher scan speeds (20 instead of 15mm/s). During the last two (10,11) passes the
laser is moved with a linear speed of 25mm/s.
At the end of the embedding step, which can be repeated in order to produce
thicker embedding layers, the sample was left to cool down to ambient
temperature and then the silicone paste was removed from the substrate base
with acetone.
Subsequently, the sensor was tested to ensure functionality. In the case of a strain
gage, we measured the resistance from the sensor to the substrate and the
resistance between the contact pads of the sensor. If the resistance to the substrate
was greater than 1 MOhms and the gage resistance was close to its value before
the embedding, then there would be strong reasons to believe that the strain gage
survived the embedding process. Of course, only a full-fledged mechanical test
would give an unambiguous answer. In the case of a thermocouple, the
calibration procedure would prove the functionality of the sensor.
70
3.7 Summary
In this chapter, the fabrication process for embedded thermo-mechanical sensors
was described. Since the purpose of this work is the proof-of-feasibility, the
sensors were fabricated on “sample” 50mm-square substrates made out of 304L
stainless steel. We presented in detail the substrate preparation sequence from
the cutting of the material to polishing and cleaning, especially with respect to
the particle-ridden environment of a manufacturing facility. We concluded that
sputter-etching the substrate prior to the deposition of the dielectric thin film
would remove the topmost contaminated layers and most of the particles which
were causing the formation of pinholes.
Then, we presented reactive pulsed-dc magnetron sputtering of aluminum oxide
as the selected method to produce thick dielectric layers at a high deposition rate.
We covered the issues that surfaced during the implementation of the technique,
especially the migration of the transition voltage and flow for an aluminum
target to lower values during its useful life, and the subsequent effects, such as
the decrease of the oxygen consumption and the deposition rate. We also
presented a guiding linear law, based on empirical results, that could be used to
predict the transition values for the oxygen flow prior to the deposition.
We described the photolithography, deposition and micro-machining processes
that were employed to produce the strain gages and thermocouples from blanket
metal films and we stressed the importance of an adhesion-promoting interlayer
between the sensor and the dielectric substrate. Then we exhibited the necessity
of protective layers that would shield the thin-film structure from the hightemperature embedding process, and presented the deposition technique
(electroplating) for producing those layers.
In the final discussion which concerned the embedding stage, we explained the
need for an additional protective layer to the original copper one, and showed
how nickel served adequately as such a layer. We also described the attempts to
71
tune the embedding process so that the sensors could be embedded successfully.
The corrections that proved most critical where the aforementioned introduction
of an additional protective layer, the pre-heated substrate, and the reduction of
the heat input and heat input rate to the substrates during embedding.
72
4
Characterization
In the following section we present data with respect to properties,
characteristics, and responses of individual components of the embedded
structure.
4.1 Composition
Compositional analysis of dielectric and sensor layers is important because their
electrical behavior depends strongly on the stoichiometry. Such analysis is not
critical for the protective and embedding layers at this stage and it will be
assumed that elemental copper and nickel are deposited during the
electroplating process, and that the composition of the fused invar powder bed is
the same as the composition of the powder particles.
4.1.1 Dielectric layers
The stoichiometry of the deposited aluminum oxide films was investigated with
X-ray Photo-electron Spectroscopy (XPS) and Electron Probe Micro-Analysis
(EPMA). The films used in both techniques had a thickness in the 3µm range.
73
CHAPTER 4: CHARACTERIZATION
XPS measurements
Even though XPS is not suitable for the exact quantitative determination of the
composition, it can be used to compare the unknown composition of a sample
with that of a “standard” sample. For example, in our case a sapphire piece
(stoichiometric, crystalline Al2O3) was used as a reference sample. Furthermore,
since XPS is a surface analysis technique (most of the signal is generated from the
top monolayers of the sample, or within 1.5nm from the surface), it was
necessary to create a compositional depth-profile of the sample by taking
measurements at the bottom of a well created by sputter-etching.
Initially, the sapphire crystal was examined by measuring the aluminum and
oxygen atomic contents at the surface and at the end of each of 3 sputter-etching
intervals. Each interval lasted for 300s and the sputter-etching rate was estimated
at ~0.13±.04nm/s (cf. “Determination of the sputter-etching rate for aluminum
oxide” on page 75).
Then, 3 films, which were grown by reactive sputtering of an aluminum target in
different conditions, were scanned. In particular, one (F1) was deposited with the
target at the transition state, achieved by an average oxygen flow equal to
5.7sccm, while the other two (F2 and F3) were deposited with the target at the
fully “poisoned” state, caused by a high oxygen flow, equal to 9.8sccm.
The first film (F1) was a thick deposit (in the order of 3µm) and a depth-profiling
of its composition was performed in the same manner with the sapphire sample:
4 measurements of the oxygen and aluminum atomic content, with a 300s
sputter-etching interval between successive measurements. Films F2 and F3 were
relatively thin (in the order of 50nm) and the sputter-etching intervals were
much sorter in duration (10, 30, and 80s). All results, in the form of oxygen-toaluminum atomic ratios are depicted in Figure 4.1.
The immediate observation is the initial sharp decrease of the oxygen content
with depth for all samples. Film F3 was analyzed in order to provide a minimum
74
scale for the observation of the reduction. Indeed, a mere 10s of sputter-etching is
sufficient to remove the layers responsible for the increased oxygen content
which is measured at the surface. In such a short interval only a small number of
monolayers can be removed and thus the excess oxygen can be attributed to the
water molecules attached to the surface after its exposure to atmospheric
humidity. Following the removal of these molecules the oxygen content remains
relatively constant for the rest of the depth-profiling analysis.
Another observation is that all sputter-deposited films show a lower O/Al ratio
than the sapphire sample. The deviation from the mean sapphire value ranges
from ~13% for film F1 to ~3% for film F3. Therefore, deposition from a fully
oxidized target surface results in a composition that is closer to the correct
stoichiometry, at the expense, of course, of the deposition rate.
3.0
Sapphire
F1 (Q Ox=5.7 sccm)
F2 (Q Ox=9.8 sccm)
2.5
F3 (Q Ox=9.8 sccm)
O/Al
2.0
1.5
1.0
0.5
0.0
0
200
400
600
800
1000
Etching time (s)
FIGURE 4.1: Oxygen-to-aluminum ratio vs. sputter-etching time.
XPS data from three samples of reactively sputter-deposited Al2O3 and one sample of
sapphire.
Determination of the sputter-etching rate for aluminum oxide
In order to determine the erosion rate of the deposited aluminum oxide by the
sputter-etching action of argon anions we used film F2 which had a measured
thickness equal to 45±15nm, as measured by stylus profilometry. The particular
75
film was grown on a polished 304L stainless steel surface with an average
roughness equal to 10 nm. Our goal was to create the compositional depth-profile
of the film by monitoring its aluminum content after short sputter-etching
intervals.
The results are presented in Figure 4.2, where it can be seen that after 350s of
sputter-etching the aluminum content has fallen to 37% (e-1) of its maximum
value. It is reasonable to argue that at this time the film/substrate interface has
been reached and therefore to calculate the sputter-etching rate by dividing the
known film thickness with the total sputter-etching time up to the critical point.
This calculation yields a rate equal to 0.13±0.04nm/s. Alternatively, if the
interface is assumed to be reached at about 300s, where the last plateau value of
atomic content is obtained, the calculated rate is 0.15±0.05nm/s.
100
Al (at. %)
80
60
40
20
Cmax/e
0
0
100
200
300
400
Etching time (s)
FIGURE 4.2: Aluminum content of deposited Al2O3 film vs. sputter-etching time.
Film F2 from Figure 4.1. Film thickness: 45±15nm. Highlighted time corresponds to
atomic content equal to 37% of maximum value.
EPMA measurements
As opposed to XPS, EPMA can give very accurate results for films of adequate
thickness (larger than 1µm). This technique gave measurements of a near-perfect
stoichiometry with an oxygen-to-aluminum ratio of 1.499 ± 0.005.
76
Perfect stoichiometry is guaranteed at high oxygen flows (Q≥Q’, cf. Table 3.7 on
page 40). Then, the target is fully covered by an aluminum oxide film, thin
enough to be sputtered even in the pure-dc mode. This film is stoichiometric and
therefore the deposit is stoichiometric, too.
Arguably, there can be inclusion of free oxygen ions or atoms in the growing film,
even at the absence of active aluminum atoms. However, judging from the
relative abundance of argon (~5:1 for a fully “poisoned” target, if all the oxygen
molecules are dissociated) and the fact that its content in the film never exceeds
4%, it is difficult to get deposits that are significantly oxygen-rich, unless the
sticking coefficient of oxygen on aluminum oxide surfaces is considerably higher
than that of argon.
Of course, it is very easy to get aluminum-rich films when the oxygen flow is
lower than QT. The target surface is only partially oxidized and the sputtered
species include both aluminum and aluminum oxide particles. Studies have
shown that it is only at conditions of a fully “poisoned” target that Al2O3 does
form on the surface of the substrate [stir71, gora78]. Oxygen-deficient films will
have a dark brown, or even black, color instead of being transparent, as the extra
aluminum atoms provide energy states within the energy gap of the – flawed –
insulator which absorb in the visible spectrum and colorize the film.
4.1.2 Sensor layers
Compositional information for the sensor layers is very significant for their
adequate characterization. The sensor layers for this work are always metal
alloys and their electrical response to strain and temperature is dependent on the
relative atomic contents of the various constituent elements.
As explained before (cf. “Deposition” on page 55), the sensing films are sputterdeposited from alloy targets. These targets were machined from bulk “blanks”
which were supplied by manufacturers of commercially available sensors. Due to
77
the different sputtering rates of the constituent elements, the alloyed targets form
a thin surface layer whose composition is such that the sputtered species get
ejected in direct proportion to their concentrations in the bulk material. This
results in a deposited film that has the same composition with the bulk target (cf.
“Composition” on page 16.)
Our goal was to determine the composition of the deposited films, and compare
the results with the nominal values available in the literature. The latter are
summarized in Table 4.1 for constantan, in Table 4.2 for chromel, and in Table 4.3
for alumel.
TABLE 4.1: Nominal composition of constantan, measured values for deposit.
Element
Nominal
(at. %)
Deposit
(at. %)
Cu
50
50.4
Ni
50
49.6
TABLE 4.2: Nominal composition of chromel, measured values for deposit.
Element
Nominal
(at. %)
Deposit
(at. %)
Cr
10
12
Ni
90
88
TABLE 4.3: Nominal composition of alumel, measured values for deposit.
Element
Nominal
(at. %)
Deposit
(at. %)
Al
2
2.5
Mn
3
1.8
Ni
94
95.7
Si
1
78
All measurements were performed with EPMA. The films were grown on single
crystal Si wafers, thus the inability to provide reliable data for the Si content in
the alumel film.
4.2 Microstructure
The microstructural characteristics of the thin films are defining to a large extent
their chemical and transport properties. However, the scope of this work is
focused more on the development of a working prototype and less on the
thorough characterization or specific property enhancement of the constituent
layers. Therefore, only simple measurements and educated estimates were
performed.
4.2.1 Dielectric layers
The decisive factors for the microstructure of sputter-deposited thin films are the
working pressure and the homologous temperature (the ratio of the substrate
temperature and the melting point of the deposit, in K) [ohri92, smit95]. The
substrate was never externally heated during sputter-deposition. The
temperature at the back side of the substrate (facing away from the magnetron
gun) was reaching 80-90°C during deposition. Considering that it was supported
only by teflon spacers, it may be reasonable to assign a temperature of ~100°C
(~373K) at the front of the substrate1. Such a temperature is typical of magnetron
sputtering systems, were substrates are relatively cooler than in simple diode
sputtering chambers [wasa92].
1. A rough, steady-state heat transfer calculation, where all the heat flux to
a 100cm2, 3mm-thick, stainless steel substrate comes from the impingement of electrons accelerated to ~300eV – assuming no energy loss due
to collisions – making up an electric current of 0.8A, yields a ~15°C difference between the front and the back sides.
79
The melting point of aluminum oxide is 2323K [incr90], and therefore the ratio
Tsub /Tm is ~0.16. Such a ratio, according to the well-known Zone Structure for
sputter-deposited coatings [thor77], can only facilitate small fibrous grains with
voided boundaries at pressures of ~1Pa. However, the substrate temperature is
far below the crystallization temperature of Al2O3, which is 1003K (730°C), and
therefore the deposit will have to be amorphous [wasa92].
Images of the deposited film, produced by Scanning Electron Microscopy (SEM),
show a glassy phase. A representative picture is shown in Figure 4.3. The domeshaped structures at the surface of the film may be either evidence of columnar
grains or features due to the surface topography of the substrate surface.
FIGURE 4.3: SEM image of amorphous Al2O3 film on a stainless steel substrate.
In order to substantiate the SEM evidence of the microscopy, Al2O3 films
deposited on 304L stainless steel were examined in an X-Ray Diffractometer. The
XRD spectra of the film and the uncoated substrate were superimposed
(Figure 4.4) so that any crystalline alumina peaks be immediately recognized.
As it is easily seen from Figure 4.4, the peaks of the aluminum oxide film
coincide with the (111), (200), and (220) peaks of the austenitic 304L stainless steel
[iccd90] and only the amorphous peak is pertinent to the film. These results are
at complete agreement with previously published work on the effect of the
substrate temperature on the microstructure of reactively sputter-deposited
80
10000
Al2O 3
304L
Amorphous
'peak'
Counts
1000
100
111 200
10
20
40
2è
220
60
80
FIGURE 4.4: XRD spectra of Al2O3 on 304L, and uncoated 304L substrate.
alumina films. In particular, no γ-Al2O3 or α-Al2O3 was detected at substrate
temperatures below 330°C [zywi96].
The microcrystalline films, especially the thermodynamically stable α-Al2O3
form, have very low thermal conductivity, high wear resistance, high hardness,
and excellent acid resistance. However, they require substrate temperatures
above 1000°C, thereby excluding stainless and tool steels as possible substrates.
There is an ongoing effort to produce microcrystalline alumina films at lower
substrate temperatures (500-700°C), using high-power (26W/cm2) pulse
magnetron reactive sputtering [schi93].
4.2.2 Sensor films
The sensor films are expected to have the characteristic columnar structure of
films which are sputter-deposited at low substrate temperatures (Tsub/Tm<0.3 for
the nickel based alloys). A limiting factor for the grain size is the very small
thickness of the films (~600nm).
4.2.3 Protective layers
The grain size of the protective layers depends solely on the method of
deposition. The sputter-deposited 2µm-thick “seed” copper layer has a fine-grain
81
structure, while the electroplated nickel layer has grains in the order of 30-50µm.
The two optical micrographs in Figure 4.5 attempt to show the dramatic
difference between the two layers.
FIGURE 4.5: Grain sizes in sputtered and electroplated copper layers.
The electroplated layer develops a distinctive elongated grain texture through its
thickness. An example of the texture can be seen in Figure 4.6.
FIGURE 4.6: Copper grains in electroplated layer.
82
4.3 Electrical measurements
The electrical characteristics of the thin-film components in the embedded
structure offer a measure for the effectiveness of these components as dielectric
layers or sensor devices. The measured quantities will be compared with the bulk
values as well as published thin-film data.
4.3.1 Dielectric strength of dielectric layers
The successful deposition of adequately insulating aluminum oxide films on
stainless steel substrate has been the most difficult step throughout the
development of the technology outlined in the “Fabrication” chapter. A large
number of sensors deposited on the Al2O3 layers was exhibiting a very small
resistance to the substrate. Since the compositional measurements ruled out the
possibility of non-stoichiometric aluminum oxide as the cause of the low
resistance paths, our attention was turned to pinholes.
Pinholes were a severe problem in the IC industry in the 60s and 70s and this
pronlem was only resolved by enforcing strict particle control in the fabrication
facilities. The concept of “clean rooms” is the result of that strategy. Pinholes are
thought to be caused by particles which are electrostatically attached to the
substrate and are incompletely covered by a blanket layer. During subsequent
processing (e.g.: rinsing) those particles might be removed, leaving a tiny hole in
their place. Therefore, any over-deposited layer will make contact with the
original substrate, and, if that layer is conductive, the contact will lead to
shorting.
In order to verify the existence of pinholes two simple methods were employed.
Initially, a small voltage (~5-10V) was applied between a “shorted” sensor and
the stainless steel substrate. After the application of the voltage, the sensor was
rendered insulated from the substrate, the resistance between them becoming
higher that 20MΩ. This method is referred to as self-healing dielectric breakdown
83
[kern73]. The surface of the sensor was then examined by an optical microscope
to reveal any damage.Figure 4.7 shows an optical micrograph of the region were
a crater was created at a pinhole site. As the voltage is applied across the
constantan/304L contact in the pinhole, a high current has to pass through a
small volume of material and generates a large amount of heat which explosively
evaporates the surrounding materials. This action eliminates the short-circuit
between the conducting layer and the metallic substrate which was caused by the
pinhole. It has to be noted that 5-10V could cause no dielectric breakdown
between a well insulated conducting film and the substrate, with a dielectric film
of the usual 3-4µm thickness.
FIGURE 4.7: Optical micrograph of damage caused by self-healing dielectric
breakdown.
A direct observation of pinholes was possible with the large depth-of-field of an
SEM. An image of two pinholes, and possibly a speck of dust which is semiburied in the sputter-deposited Al2O3 layer is shown in Figure 4.8. below.
In order to develop a measure of the dielectric strength of the deposited
aluminum oxide films, metal conductors (aluminum and constantan) were
deposited on a 3µm-thick oxide which was sputter-deposited on polished
84
FIGURE 4.8: SEM image of Al2O3 film with pinholes.
stainless steel substrate. The conductors were photolithographically shaped in
200µm squares. The thickness of the dielectric was measured by a step
profilometer. Then, using a dc power supply, a voltage was applied between the
conducting patches and the substrate with the aid of copper electrodes. The
voltage varied continuously from 5 to 32V. In all cases, the dielectric broke down
at 30±1V, thereby exhibiting a dielectric strength of ~10V/µm, which is near the
low end of the reported values [chen97]. Yet, that value was adequate for the
purposes of this work. It is expected that microcrystalline aluminum oxide, and
especially the stable α-phase, would show higher dielectric strength values
[dörr84].
4.3.2 Resistivity of sensor materials
The resistivity of the sensor films was measured with a standard 4-point probe
station. In its simplest incarnation the probe consists of four equidistant tips
which are pressed on to the film surface. The four tips lie on a straight line and
the outer two supply a known constant current I while the inner two measure
voltage V. As long as the film thickness is a lot smaller that the distance between
the tips the sheet resistance RS (in Ω/")of the film can be calculated by Eq.(4.1)
85
[smit95].The resistivity ρ of the material is given by multiplying the sheet
resistance with the thickness of the film.
π V
R S = -------- ---ln 2 I
(4.1)
We prepared three films, one for each of the three sensor materials, and present
the results along with data from the literature inTable 4.4.
TABLE 4.4: Resistivity and thickness of alumel, chromel, and constantan films.
Material
Published ρ
(Ω.nm)
Measured ρ
(Ω.nm)
Thickness
(nm)
Alumel
260a
440
630
Chromel
594a
819
630
Constantan
490b
840
1050
a. Measured for wires with 50.8µm diameter [alda84].
b. Bulk value at 25°C [crc85].
The considerably higher resistivity of the deposited films (70% higher than the
reported bulk values for alumel and constantan) can be explained by the nature
of the films grown by sputtering at low homologous temperatures. Indeed, all
the experimental sensors were deposited without any thermal activation of the
substrate so that the surface mobility of the physisorbed species on the substrate
was rather low. In such a case, the resulting microstructure is strongly
microcrystalline and at best columnar, in which case the largest dimension of the
grains is equal to the film thickness (a few hundred nm in our case). The small
grain size results in a large number of grain boundaries which act as scattering
centers for the electrons and therefore increase resistivity.
86
4.4 Strain-gage characterization
The investigation of a strain-gage response to applied loads is of central
importance to the characterization of the embedded system. It is true that a
sensor mounted on the surface of a component will bear higher strains than in
any other position, therefore we chose to test the strain gages in their most
vulnerable state. This method would also be revealing for the mechanical
integrity of the dielectric films under strain.
4.4.1 Experimental setup
The simplest configuration for such a test is the four-point bend test (4PBT) of a
beam with a strain gage deposited on its surface. This setup allows an exact
comparison of the strain gage output with a calibrated commercial gage and
beam-theory calculations as well. The latter comparison is possible because the
bending moment is constant between the two inner pins (cf. Figure 4.9), and
therefore the resulting outer-fiber strain is constant in the same region.
The beam substrate was of a rectangular cross-section and its dimensions are
given in Table 4.5.
TABLE 4.5: Dimensions and moment of inertia for beam substrate.
Length
(10-3 m)
Width
(10-3 m)
Height
(10-3 m)
Moment of inertia
(10-12 m4)
50
23.4
2.96
50.55
The substrate was cut out from a 304L stainless steel plate. The nominal Young’s
modulus and yield strength for this material are 193GPa and 170MPa [asm94]
respectively, at room temperature.
Following the procedures described in “Substrate preparation” on page 28, a
3µm-thick aluminum oxide layer was sputter-deposited on the polished surface
of the substrate. No shadow mask was used and the layer covered the surface in
87
its entirety. Then, a strain gage was deposited from a constantan target according
to the processes described in “Sensor layer” on page 53. Its longest dimension,
along the longest substrate dimension, was 8mm and the line-width was 75µm.
The strain-gage resistance was measured equal to 225Ω.
A calibrated constantan gage was attached next to the experimental one. The
commercial gage was a thin-film sensor deposited on a polyimide foil carrier.
This carrier was bonded to the aluminum oxide film on the substrate surface by
means of a fast curing adhesion agent. The commercial gage resistance was 120Ω.
Both strain gages were connected to bridge circuits prior to the mounting of the
substrate on the 4PBT apparatus. Connections to the sensor leads were made by
using regular solder. The commercial gage was connected to a specialized strain
indicator box (Model P-3500 from the Measurements Group) with the ability to
provide both a digital readout of the strain as measured by the gage and an
output signal to be processed by external Data Acquisition Systems (DAqS). The
experimental gage was connected to a bridge circuit which was built in the
laboratory and was controlled by the testing equipment.
The testing equipment was manufactured by MTS and could be equipped with a
2240N load cell. The accompanying control electronics could be used to operate
the system in displacement-control or load-control mode. The substrate was
mounted on a 4PBT rig in such a way that, when load was applied, the sensors
would be in tension. The dimensional details of the setup are depicted in
Figure 4.9.
Also, the substrate had to be fully insulated from the testing equipment. In order
to achieve this, the back surface (upper side in Figure 4.9) was covered by Kapton
tape and all wire connections were wrapped with insulating tape. Inadequate
electrical insulation would often result in erratic sensor behavior, especially for
sensors that featured a finite (~1MΩ) “shunt” resistance to the substrate.
88
Strain gages
10 mm
20 mm
40 mm
FIGURE 4.9: Dimension details of the 4-point bend test setup.
Data processing was facilitated by a PC equipped with data acquisition
equipment controlled by a LabView application. The 4 voltage signals read by
the DAqS were generated by the applied load from the load cell, the cross-head
displacement from the ram, the bridge circuit with the experimental sensor, and
the bridge circuit with the commercial sensor, as amplified by the strain indicator
box. Finally, a common ground was used for all testing and control equipment.
4.4.2 Experimental objective and results
Clearly, one of the most important requirements for a strain sensor is the absence
of any hysteresis effects when the tested structure is experiencing elastic strains
only. Furthermore, for the particular implementation of our design it is important
that the aluminum oxide films that encapsulate the sensor are not strained to
fracture while the substrate is still behaving elastically. These two requirements,
no history effects and no “carrier” failure in the elastic regime of the metal
matrix, dictated the form of the load curves that were imposed on the tested
substrate.
89
All experiments were performed in the displacement-control mode so that timedependent localized deformation could be easily detected, and extended yielding
could be prevented. In the load-control mode the machine moves the ramp until
the pre-defined load level is achieved. Apparently, if the applied load causes the
stress to exceed the yield limit, the resulting deformation will be rather extensive.
During the experiments the ram displacement was increased or decreased until a
pre-defined load level was reached. The system was kept at that displacement for
a short period of time (one or two minutes) and then the displacement was readjusted so that the next load level would be reached.
Elastic regime
In the initial experiments the maximum applied stress was kept below the
nominal yield strength of the stainless steel substrate (170 MPa). In Figure 4.10
we present the stress at the outer fibers of the test sample during the 4PBT, as
calculated by beam theory based on the applied load, and the voltage output of
the bridge that contained the experimental – deposited – sensor. Both quantities
are plotted against time.
-15
Experimental gage
200
-10
150
-5
100
0
stress (MPa)
50
0
5
0
200
400
600
800
1000
Experimental gage output (Volts)
Outer fiber stress (MPa)
250
10
Time (s)
FIGURE 4.10: Outer fiber stress and of deposited strain gage output from 4-point bend
test.
Test run in displacement control mode. Strain gage in tension.
10% of total data shown for clarity.
90
The form of the loading and unloading curve permits the comparison of the gage
output for load levels of the same magnitude which were reached in different
manners (i.e. after increasing and after decreasing the load). This comparison for
loads resulting to stresses of ~25, ~50, ~80, and ~100 MPa, shows that the gage
output levels remain constant irrespective of the loading history. This is a first
evidence that, in the purely elastic regime, there are no history effects in the
behavior of the deposited sensor.
It can also be observed that while the displacement was kept at a constant level,
the load (and the resulting stress and strain at the outer fibers) would relax with
time, even though the maximum stress was below the yield strength of the
material. This time-dependent behavior was attributed to the local deformation
of the polymer tape that had been attached to the back of the substrate (top
surface in Figure 4.9), as the hardened-steel pins would cause the underlying
tape to “flow”. A supporting argument for this conclusion is the much smaller
scale of the effect when the displacement has been reduced from a higher value
(cf. Figure 4.11.b), rather than increased from a lower one (cf. Figure 4.11.a).
(a)
(b)
-20
85
85
-20
Displacement (x10 µm)
-15
75
80
-15
75
Stress
80
Displacement
Displacement (x10 µm)
Displacement
Stress
-10
480
520
560
Time (s)
70
600
-10
720
740
760
70
780
Time (s)
FIGURE 4.11: Displacement and calculated outer fiber stress as functions of time.
a) Measurement after displacement (absolute value) increase,
b) Measurement after displacement (absolute value) reduction.
The time scale is the same as in Figure 4.10. 10% of total data shown for clarity.
91
However, the litmus test for the correct functionality of the experimental gage is
the direct comparison of its output with that of the calibrated commercial strain
sensor. The graphs in Figure 4.12 depict the experimental gage output as a
function of the commercial gage reading, and the output of both sensors as
functions of the calculated outer fiber strain. The latter was obtained by applying
beam-theory calculations to the sample, using the measured load values as input.
(a)
50
100
(b)
150
-6
ε (x10 ) = a + b * V
Experimental gage
600
-6
400
-4
Calibrated gage
-2
200
0
0
800
800
Nominal strain (x10-6 )
-8
0
-6
Nominal strain (x10 )
-10
600
a
b
Value
-9.315
-69.065
Error
0.28121
0.050369
400
200
Nominal fracture strain
for aluminum oxide
200
400
600
0
800
Calculated strain (x10-6)
0
0.0
-2.5
-5.0
-7.5
-10.0
FIGURE 4.12: a) Experimental gage output and nominal strain vs. calculated strain and
stress.
b) Nominal strain vs. experimental gage output.
Nominal strain as measured by calibrated gage. All other values from Figure 4.10.
10% of total data shown for clarity.
From Figure 4.12.a it is understood that the experimental gage output is perfectly
linear with the applied load and the calculated resulting stress and strain.
Moreover, there is no visible history effect even though the loading was far from
monotonic (cf. Figure 4.10). The spread of strain values measured by the
calibrated gage at a particular stress “level” is attributed to both the resolutionlimited method of acquisition and the electrical noise generated during the
amplification and transport of the signal.
92
Another interesting conclusion from the first graph is the agreement between the
calibrated gage readout and the calculated strain values. The linear fit of the
calibrated gage output values is given by:
ε
fit
calibr
–6
= – 8.98 ×10
+ 0.98 ⋅ ε calc
(4.2)
The intercept is exceedingly small in absolute numbers and one to two orders of
magnitude smaller than the elastic strains measured on the sample (50-600µε).
Also, the slope of the curve is very close to unity, even though the elastic
modulus value used in the calculations was taken from the literature. This
agreement provides confidence in the validity of the theory-based calculations.
Furthermore, the relation of linearity between the measured and calculated
values is solid evidence that the proportional elastic limit was not exceeded
during this portion of the 4PBT, as the calculated values were based on linear
elasticity theory.
The second graph (Figure 4.12.b) provides a calibration curve for the deposited
strain sensor, for the particular signal acquisition and amplification system. This
curve is also a linear fit where the relation between the measured voltage and the
corresponding strain is given by:
–6
ε calibr = – 9.31 ×10
– 69.07 ×10
–6
⋅V
(4.3)
whereas the voltage output range is 0-(-10)Volts.
The maximum uncertainty in Eq.(4.3) (as expressed by the combined errors for
the intercept and the slope coefficient) is realized at very small signals (<-0.1V)
and can be as large as 1.8-3%. At high output values (~-10V) the error can be as
small as ~0.11%. The sources of this uncertainty are again the electrical noise
generated by the strain indicator device that was attached to the calibrated gage,
and the low resolution of the acquisition equipment.
93
Inelastic regime
In the second set of experiments, the sample was subjected to loads that resulted
in an outer fiber stress greater than the nominal yield stress (170MPa) of 304L
stainless steel. All equipment remained the same except for the data acquisition
card attached to the bridge containing the experimental gage. The previous
acquisition system reached its maximum output during loading in the elastic
regime and was replaced by a card with lower amplification. In Figure 4.13.a we
present the voltage output of the bridge with experimental gage and the
calculated outer fiber stress as functions of time. In graph (b) of the same figure
we present the nominal strain as measured by the calibrated strain sensor.
(a)
(b)
500
300
-2
200
-1
100
0
stress (Mpa)
1200
1400
Time (s)
1600
1
Outer fiber stress (Mpa)
-3
Nominal strain
1400
400
700
300
0
200
100
0
-6
Calibrated gage (x10 )
400
Experimental gage
0
500
-4
-700
stress (Mpa)
1200
1400
1600
Time (s)
FIGURE 4.13: a) Calculated outer fiber stress and output of deposited strain gage.
b) Calculated outer fiber stress and strain measured by calibrated strain gage.
From 4-point bend test in displacement control mode. Strain gages in tension. Inelastic
regime.
The dashed lines in the graphs have been included to emphasize the fact that
while the applied stress returns to the same level, the strain (as measured by both
sensors) does not. Such behavior is strong evidence of plastic deformation. A
better visualization of the permanent deformation can be given by the stressstrain curves based on the data enclosed in the rectangles in Figure 4.13.
In the two graphs of Figure 4.14 the outer fiber stress is plotted against the
output of the strain gages. The direction of loading is indicated by different
94
markers and the stress values lower than the yield strength of the substrate
(170MPa) are linearly fit.
(b)
(a)
a
b
200
150
inc
=a+b*V
σ
-142.01
-114.53
σ
Increasing load
Decreasing load
"Plastic" load
100
σ
dec
=a+b*V
a
-161.53
b
-114.07
inc
a
b
200
=a+b* ε
-4.2549
0.20066
150
Increasing load
Decreasing load
"Plastic" load
100
σ
dec
a
b
50
50
-1.5
-2.0
-2.5
-3.0
-3.5
=a+b* ε
300
600
-21.69
0.19757
900
1200
-6
FIGURE 4.14: a) Outer fiber stress vs. experimental gage output.
b) Outer fiber stress vs. nominal strain.
Linear fit parameters for elastic stress values (<170MPa). Data from Figure 4.13
(within rectangles).
Both strain gages register a permanent strain, roughly equal to 0.01% (as
measured by the calibrated sensor). Also, the loading and unloading curves have
almost identical slope coefficients (within 0.4% for the experimental sensor and
within 1.5% for the calibrated sensor). Since the slope coefficient is the apparent
elastic modulus of the system this agreement further enhances the credibility of
the results.
The data corresponding to the first loading-unloading sequence in Figure 4.13
were not included in the calculation of the slope coefficients due to the presence
of a small history effect. During that sequence, the maximum applied stress is
just below (~165MPa) the nominal yield strength of the stainless steel substrate.
It is noted that during testing in the elastic regime the outer fiber stress did not
exceed 135MPa and no history effects were observed.
95
A closer look at the response of the sensors prior to the subjection of the substrate
to “plastic” loads is attempted in Figure 4.16. For reference, the actual loading
history is shown in Figure 4.15 with identifiers for the segments of interest.
800
-4
-3
600
-2
400
-1
α
γ
β
0
200
1
Experimental gage
Stress
0
1100
1200
1300
1400
1500
2
1600
Time (s)
FIGURE 4.15: Maximum stress and experimental gage response prior to plastic
deformation.
Data from Figure 4.13.a, during elastic loading and unloading (segments
(a)
180
-122.96
-107.79
α
a
b
160
180
α) σ = a + b * V
β) σ = a + b * V
140
a
b
β
-143.85
-115.35
γ) σ = a + b * V
120
a
b
100
-142.03
-114.53
γ
80
60
160
(b)
α) σ = a + b * ε
a
b
120
100
α
5.1923
0.18978
β) σ = a + b * ε
140
α, β, γ).
a
-4.1686
b
0.19931
β
γ) σ = a + b * ε
a
b
-4.2525
0.20066
γ
80
60
40
-1.6
-1.8
-2.0
-2.2
-2.4
Experimental gage
-2.6
-2.8
40
200
300
400
500
600
700
800
900
Nominal strain
FIGURE 4.16: a) Stress-strain curves for the experimental sensor.
b) Stress-strain curves for the commercial sensor.
All data from segments α, β, and γ in Figure 4.15.
Notice that segment γ is the “increasing load” segment in Figure 4.14
96
Again, the calibrated sensor response is juxtaposed to that of the deposited strain
gage in order to show that the history effect detected by the latter is not an
artifact of the sensor. This history effect manifests itself not so much as a
permanent strain after a “premature” plastic deformation, as a change in the
apparent stiffness of the system (4.8-5.4% as indicated by the calibrated sensor).
Since it was observed that higher loads, which did result in plastic deformation,
did not alter the system’s elastic response, it can be assumed that the particular
change in Figure 4.16 was due to a “re-alignment” effect which was triggered by
the particular load.
Finally, in analogy to the elastic regime case, we can attempt a calibration of the
sensor with respect to the acquisition system used in this case with the higher
loads. Indeed, fitting the nominal strain, as measured by the calibrated sensor, to
the output of the deposited strain gage yields a straight line (see Figure 4.17)
which is described by:
–6
ε calibr = – 680.27 ×10
– 567.66 ×10
–6
⋅V
(4.4)
1400
-6
1200
1000
800
600
-6
ε (x10 ) = a + b * V
400
200
0
-1.0
-1.5
-2.0
Value
Error
a
b
-680.27
-567.66
0.90623
0.39461
R
0.9987
NA
-2.5
-3.0
-3.5
FIGURE 4.17: Nominal strain vs. deposited sensor output.
Calibration curve based on all data from Figure 4.13. 10% of total data shown for
clarity.
97
The voltage output range from the bridge containing the experimental sensor is
0-(-10) Volts.
It must be noted that the signal data from the two sensors remain in very good
agreement through the elastic and plastic sections of the experiment, but the
intercept is quite large (-680.27µε) and corresponds to a compressive strain. The
explanation of this discrepancy lies with the lack of data from the deposited gage
with the lower amplification system near the zero-load region, before the onset of
plastic deformation in the substrate.
4.5 Summary
In this chapter the measurements regarding the chemical composition of the
dielectric and sensor films, as well as resistivity measurements for the latter were
presented. Also, the experimental results from 4-Point Bend Tests (4PBT) of a
deposited sensor are included and commented in their correlation with stress
and strain values acquired with calibrated media.
The aluminum oxide dielectric films were found to be almost perfectly
stoichiometric (O/Al=1.499±0.005) when measured by the more accurate
Electron Micro-Probe, and slightly aluminum rich when measured by X-ray
Photoelectron Spectroscopy (XPS). XPS measurements revealed a dependence of
the composition on the oxidation degree of the aluminum target during reactive
sputtering. In particular, high oxygen flows that bring the target to the fully
“poisoned” state will result in compositions closer to the stoichiometry, at the
expense of the deposition rate. The measured resistivity of the deposited sensor
films was significantly (up to 70%) higher than the bulk values reported in the
literature. This deviation was attributed to the increased electron scattering from
the proliferation of grain boundaries in thin films which were sputter-deposited
at room temperature.
98
A 304L stainless steel substrate was subjected to two tests in the 4-Point Bend
geometry. We deposited a constantan strain sensor on the surface where tensile
stress would be observed, and next to it we attached a calibrated strain gage. In
the first test the stress at the outer fibers of the substrate did not exceed the
nominal yield strength of the 304L stainless steel. The response of both sensors
was linear with respect to the stress and strain values that were obtained from
beam theory calculations. The measured dimensions of the substrate as well as
the load values from the testing equipment were used in the stress calculations.
The calculated strain was obtained using a nominal elastic modulus for stainless
steel. The correlation between measured outputs and calculated values was very
good. The ratio between the calibrated strain gage values and the calculated
strain was 0.98. By plotting the calibrated gage output against the voltage
measured by a bridge circuit which contained the deposited sensor, we were able
to obtain a linear calibration curve for the experimental sensor. The curve is
specific to the amplification system attached to the experimental gage.
In the next 4PBT the applied loads were raised in order to cause stresses that
would exceed the yield stress of the substrate. Plastic deformation was registered
by both sensors upon unloading. The stress-strain slopes of the loading and
unloading curves (corresponding to elastic applied stresses) were equal to within
0.4% and 1.5% for the experimental and the commercial gage respectively. The
resulting permanent strain was ~0.01% as measured by the calibrated gage.
Finally, throughout both the elastic and plastic stages of the second test the
correlation between the commercial and the experimental gage remained very
good, thus enabling the calculation of new calibration curve for the experimental
sensor. The linear curve was specific to the low-resolution signal amplification
system used in this test.
99
5
Modeling
The subject of this chapter is the effort to model the embedding process
employing a finite element analysis software package (Abaqus). The first goal of
this model is to predict the temperature at the copper/stainless steel interface by
simulating the heat transfer during the laser deposition process. The primary
purpose is to predict the optimum thickness of the protective layers in order to
minimize their deposition time while still adequately shielding the thin films
from the high temperatures produced during embedding. Ultimately, a refined
model can be used to predict the mechanical behavior of the substrate during
and after the laser deposition process.
5.1 Model Description
5.1.1 Geometry
The model attempts to replicate the geometry and the dimensions of the actual
substrate, the protective, and the embedding layers. For simplicity the model is
comprised of 4 strata. The first stratum is the stainless steel substrate and is a
50x50x3mm rectangular slab. The second stratum contains the protective copper
layer as well as parts of the nickel layer and the powder bed. The third stratum
contains the rest of the nickel layer (whatever is deposited on top of the copper
100
CHAPTER 5: MODELING
layer) and part of the powder bed, while the fourth stratum contains only
powder. Each of strata 2-4 have a thickness of 1mm and their origins are offset by
20mm from the origin of the first stratum, along the Y-axis, in order to account
for the part of the stainless steel surface that is left uncovered to provide access to
the sensor leads. The structure is depicted in Figure 5.1 on page 102.
Mesh Selection
Since the main goal was to calculate the temperature at the interface between the
copper and the stainless steel layer, a high nodal density along the Z-axis was
deemed necessary.
The nodal densities along the two principal axes in the horizontal plane were
selected in order to facilitate the simulation of the moving laser beam. The
diameter of the laser beam spot on the part is 2.5mm with a cross-sectional area
of 4.91mm2. The initial nodal densities along the three axes are mentioned in
Table 5.1. The particular values lead to the construction of brick-like, 3D elements
with a top face area equal to 5mm2 (2mm x 2.5mm). This value is only 1.8%
larger than the actual cross-sectional area of the laser beam spot.
This approximation allows the assignment of a maximum heat input value to an
element face which results into a heat flux equal to the actual value imparted by
the laser beam. The immediate benefit is the simplification of the simulation of
the moving heat source.
Unfortunately, the elements that are generated by the mesh according to Table 5.1
are quite “flattened”, and the result is a non-physical fluctuation of the
temperature at certain nodes during the heat transfer calculations. The
fluctuations usually prompted the analysis to be aborted.
In order to determine the cause of the problem two simple models were built
(each with 4 strata of equal dimensions), using a single material for each stratum
to avoid geometry effects. The first model used the nodal densities in Table 5.1.
101
CHAPTER 5: MODELING
30 mm
30 mm
Invar
45 mm
24 mm
30 mm
Invar
Nickel
50 mm
40 mm
30 mm
Invar
Nickel
Copper
20 mm
50 mm
50 mm
Invar powder
304L stainless steel
55 mm
FIGURE 5.1: Materials and dimensions of the 4 strata in the geometric model.
102
CHAPTER 5: MODELING
TABLE 5.1: Nodal density along the three axes. Initial model.
Nodal density
(nodes/mm)
X-axis
Y-axis
Z-axis
0.4
0.5
2
The second model used values increased by 5-fold along the horizontal axes. In
this model the horizontal dimensions were reduced by a factor of five so that the
elements were nearly cubic, but the heat flux remained at the same level. The
behavior of the second model showed no evidence of temperature fluctuations
and its analysis completed succesffuly. Therefore, it was deduced that the
densification of the mesh would be necessary for the successful completion of the
analysis
In order to retain a relatively small number of nodes and the approximation of
the laser beam spot with a superset of element faces, the mesh was densified by
increasing 2-fold the horizontal nodal densities (Table 5.2).
TABLE 5.2: Nodal density along the three axes. Final model.
Nodal density
(nodes/mm)
X-axis
Y-axis
Z-axis
0.8
1
2
This quadrupled the total number of nodes in the model to 21,443, without
including the nodes in the surrounding material which provides for more
realistic boundary conditions (cf. “Boundary and initial conditions” on page 107,
and “Invar powder” regions in Figure 5.1). At the same time, the elements at the
top-most layer were grouped in element sets. Each new element set consists of
four 1mm x 1.25mm elements which share an edge along the Z-axis, and has a
top surface area of 5mm2. Therefore, it is identical in size to the elements
produced by the mesh in Table 5.1
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CHAPTER 5: MODELING
Simplifications
This simple model does not take into account the thin-film multilayer which is
located between the stainless steel substrate and the copper protective layer. The
multilayer consists of a 10-20nm titanium adhesion layer, two 3-4µm-thick
aluminum oxide films and a second 10-20nm titanium adhesion layer (on top of
which a copper “seed” layer is sputter-deposited in order to permit the growth of
the copper protective layer by electroplating), thus its total thickness is roughly
6-8µm. The sensor circuit which is enclosed in the dielectric envelope has a very
small area compared to the envelope itself (6-7%) and its thermal properties do
not affect significantly the thermal behavior of the multilayer.
5.1.2 Material Properties
Calculation of the temperature profile in a heat transfer problem requires the
knowledge of the thermal properties for all the materials involved. In the actual
object the constituent materials are: solid 304L stainless steel, solid copper, solid
nickel, invar (Fe-36Ni) in both solid and powder form, and finally the titanium
and aluminum oxide thin films.
As explained before, the thin films are not included in this heat transfer
calculation. The thermal properties required by the FE model include specific
heat, thermal conductivity, latent heat, and liquidus and solidus temperatures.
Finally, the density of each material is also required. The arithmetic values for the
aforementioned properties are reported in Table 5.3.
Latent Heat
The inclusion of the latent heat effect in the analysis makes the problem nonlinear and requires the use of first-order 3D elements [hks97]. Such elements
have 8 nodes and 6 faces instead of the 20 nodes and 6 faces in second-order
elements. In first-order elements the temperature is linearly interpolated between
adjacent nodes.
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CHAPTER 5: MODELING
TABLE 5.3: Density and thermal properties of materials in the FE model.
Material
Density
(kgm-3)
Specific
heat
(JK-1)
Thermal
conductivity
(Wm-1 K-1)
304L
7900a
477a
14.9a
Copper
8933b
385b
401b
Nickel
8900b
444b
90.7b
Invar
8000c
515d
11
a.
b.
c.
d.
e.
f.
Latent heat
(JK-1)
Solidus/
Liquidus
(°C)
282290e
1440/1450f
Room temperature [frad93].
Room temperature [incr90].
Full theoretical density.
Room temperature [toul70].
Interpolated value from Fe and Ni data [crc85], for 64Fe-36Ni.
[asm73]
Simplifications
In reality, invar is first deposited in powder form and its volume is gradually
solidified as the laser beams scans the surface. The solidification results in a
density increase, accompanied by an increase in the apparent thermal
conductivity, which cannot be modeled with the tools available. The reason is
that the model properties cannot be altered in real time while maintaining the
temperature values achieved by the preceding analysis steps.
Therefore, the density and the thermal conductivity of the invar layers have to be
defined at constant values throughout the whole analysis. Since it is rather
difficult to estimate the apparent thermal conductivity of the powder, and the
densification to full density is instant upon the melting caused by the laser beam,
it was deemed both practical and realistic to assume a fully solid invar layer,
which is, in effect, being remelted by the laser beam. The disadvantage of this
assumption is the incorrect calculation of the lateral heat conduction in the plane
of the invar layer.
Furthermore, all thermal properties were considered as temperature
independent, which does not reflect reality accurately. In particular, the thermal
105
CHAPTER 5: MODELING
conductivity of copper decreases with increasing temperature, while its specific
heat increases. Invar and nickel, as magnetic materials, exhibit a non-monotonic
variation of their thermal conductivities and specific heats in the vicinity of their
Curie temperatures (invar: TC=279°C; nickel: TC=358°C). As temperature
decreases from the melting point (or solidus), thermal conductivity and specific
heat values decrease for both materials. Thermal conductivity exhibits a cuspshaped minimum at the Curie temperature, while specific heat reaches a
minimum above TC and then rises to a discontinuity at the magnetic transition
temperature. Below TC both properties decrease in a monotonic fashion [toul70].
Also, the specific heat at room temperature is lower than that at high
temperatures for all materials. The opposite is true for the thermal conductivity
with the exception of invar, whose thermal conductivity increases with
temperature, at least up to 540°C [toul70].
5.1.3 Simulation of a moving heat source
The actual laser beam is scanning the part surface with a speed pre-defined in the
LabView code [“Sample LabView file” on page 125] that controls the robotic arm
which carries the end of the optical fiber connected to the lasing cavity
[“Embedding Layers” on page 65]. By quantizing the powder surface in
2x2.5mm rectangles we can modulate the heat flux input in the model in a way
so that a moving heat source is simulated. This technique is outlined below.
Let’s assume a strip of the powder bed along the X-axis of length a and width b.
The strip is exactly divided in m element sets, each one of length a’ and width b,
so that a=ma’. Let’s also assume that the spot size of the beam is a’xb and the
speed of the laser beam along the same axis is ua. All these quantities are
depicted in Figure 5.2
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CHAPTER 5: MODELING
ua
element set 2
a’
element set m
element set 1
laser spot
at t=t1
b
a
FIGURE 5.2: Quantization of the powder bed in elements, top view.
In fact, the laser is already on when its spot gradually enters the powder bed,
therefore, some heat will be input to the element during a time interval of 2a’/ua
(see also Figure 5.3). The power flux to an element increases linearly from zero to
maximum (where the spot coincides with the element) during the first half of
that interval and then decreases to zero during the second half, as the laser beam
moves to the adjacent element. Due to the adjacency condition, the heat flux
increases in the second element while it decreases in the previous one, but the
total heat flux remains constant as in the actual experiment. Also, due to this
overlap the total time during which power is input to the strip is (m+1)a’/ua.
Therefore, selecting the top face size of the element set to be equal to the spot size
of the laser beam allows a simulation of the moving heat flux as it becomes
possible to prescribe a time-varying heat flux to a face of an element set and
address the overlapping heat input intervals through a universal timing system.
The heat flux for each of the first two element sets in Figure 5.2 as well as the
total flux are shown in Figure 5.3.
5.1.4 Boundary and initial conditions
It may be helpful to note at this point that, in the context of embedding, the
notion of the substrate includes the stainless steel base and the two protective
layers. During the actual embedding process the substrate is pre-heated by
contact with a copper slab sitting on hot plate (cf. page 67). The equilibrium
temperature of the copper piece is ~350°C. Our assumption is that the
107
CHAPTER 5: MODELING
heat flux
element set 1
t
heat flux
element set 2
t
heat flux
all element sets
0
t1
2a′
-------ua
t
a′
( m + 1 ) ----ua
FIGURE 5.3: Heat flux for element sets 1 and 2 in Figure 5.2 and total heat flux.
temperature at the bottom face of the substrate remains constant and equal to
350°C during the embedding process, the copper-hot plate system effectively
being a heat sink. Furthermore, the entire substrate and the powder bed are
assumed to have an initial temperature equal to 320°C.
The top surface of the powder bed (considered as fully dense solid in the model)
as well as the uncovered surface of the substrate are all considered thermally
insulated, except for the portion of the powder bed where heat flux is imparted
from the top. The heat flux value is calculated by dividing the laser power – as
measured at the exit of the laser source – with the cross-sectional area of the laser
beam. The latter has been measured equal to 4.91mm2. In all embedding
experiments the output power of the laser was equal to ~1.8kW, and therefore
the heat flux was ~360MWm-2. Allowing for a reflectivity coefficient equal to 0.1,
the value used in the calculations was 330MWm-2. Even though invar has a very
high reflectivity in the infrared, the surface of the powder bed is very rough and
inhomogenous, thus, a low reflectivity coefficient was chosen.
Prior to embedding, the deposited powder bed extends around the substrate in
order to provide lateral support to the powder placed on top of the nickel layer.
108
CHAPTER 5: MODELING
This surrounding mass of invar powder has been included in the model as a wall
with a height equal to 6mm and thickness equal to 2.5mm. The nodes that reside
on the outer boundary of the wall (“Invar powder” in Figure 5.1 on page 102) are
assumed at contact with an infinite heat reservoir held at a constant temperature
of 320°C. The thermal properties of the invar powder are taken from Table 5.3,
with two exceptions: the density is assumed to be 50% of the theoretical value
(i.e.: 4000kgm-3), and the latent heat is assumed to be zero in order to simplify
calculations.
Simplifications
Film (convection) cooling effects or radiation effects are not taken into account.
The simulation of convective cooling is beyond the scope of this simple model as
it would require the experimental determination of the convective film
coefficients for the stainless steel and the solidified invar layer. In reality,
convective cooling is enhanced by the flow of dry nitrogen towards the substrate.
The nitrogen is introduced to minimize oxidation of the powders and the
surfaces, but provides non-uniform cooling as it is injected by a permeable
shroud which is attached to the robotic arm that moves the laser over the
substrate.
Radiation effects are not taken into consideration as they become important only
at very high temperatures. The total net radiation flux emanating from an object
at an absolute temperature T can be calculated by the Stefan-Boltzmann law
[hks97]:
4
4
W (T ) = ε ⋅ σ ⋅ T – (T 0)
(5.1)
where ε is the emissivity of the object (with a value in the 0-1 range, 1 being the
emissivity of a black body), σ is the Stefan-Boltzmann constant, equal to
56.70nWm-2 K-4 [crc99], and T0 is the ambient temperature.
109
CHAPTER 5: MODELING
The emissivity of invar near its solidus temperature (1440°C) is close to 0.33
[wahl52]. As the output power flux of the laser used in the embedding process is
~360MWm-2, we realize from Eq.(5.1) that the temperature necessary for
radiation fluxes equal to 5% of the incident power flux is above 5200°C. This is
well above the vaporization temperatures of both iron (2861°C) and nickel
(2913°C) [crc85], and it can be assumed that only an exceedingly small amount of
the illuminated mass is evaporated.
5.2 Results
An actual heat transfer calculation involves a single pass of the laser beam across
the substrate, which is completed in 3.5s, and a natural cooling of the substrate
for another 16.5s (3.5s less than the 20s cooling time used in reality). The
calculation is completed in 77 hours with the software running on a 333MHz
UltraSPARC-IIi system. The complete heat transfer calculation for the 11 laser
passes would scale accordingly in time and data-space, and it was not attempted
due to practical limitations.
The first priority was to check the calculated results with previously recorded
experimental data. The latter, in Figure 5.4, were obtained with a type-K
thermocouple connected to a Data Acquisition system, polling the sensor at a
frequency of 10Hz. The thermocouple was held in contact with the nickel layer
surface while the laser was scanning the invar powder bed.
The calculated values correspond to the temperature of a node in the nickel-invar
interface, positioned at the same location with the thermocouple in the
embedding experiment. The simulated laser beam was also moved with the
speed used in the experiment (15mm/s).
It can be seen that the calculation describes adequately well the process.
However, there are four points that deserve closer attention. First, the calculation
110
CHAPTER 5: MODELING
1400
Experim.
Calculated
Temperature (°C)
1200
1000
800
600
400
200
0
85
90
95
100
Time (s)
FIGURE 5.4: Experimental and calculated temperature at a point on the Ni-invar
interface.
predicts a lower heating rate than the experiment. This can be attributed to the
fact that in reality the invar layer is initially in powder form and therefore
exhibits an apparent thermal conductivity much lower than the bulk value which
was used in the calculations.
Next, the calculated cooling rate is higher than the observed one. This
disagreement is most likely caused by the variation of the material properties
with temperature. As mentioned in page 106, the specific heat values at high
temperatures are larger than the room-temperature values used in the
calculations, and therefore the actual cooling rate could have been overestimated.
The temperature values which lie below the measured initial steady state
temperature of ~350°C can be ignore as they were caused by temporary loss of
contact with the substrate. During the deposition of subsequent invar layers the
thermocouple was fused to the substrate and no such cooling effects were
observed.
Finally, the calculations overestimate the temperature at the end of the cooling
cycle by as much as 20°C. A possible explanation for this is the disregard of the
cooling effect imparted by the nitrogen flow.
111
CHAPTER 5: MODELING
Another concern regarding the validity of the model was the maximum
temperature at the invar-nickel interface. As experiments have shown, the
interface is completely remelted (cf. Figure 3.16 on page 69), which leads to the
conclusion that the nickel is heated at least to its melting point (1453°C).
Temperature calculations for the nodes at the interface, along the center-line of
the laser path and at a path offset by 1mm from the center-line, have produced
the data in Figure 5.5.
1500
Nickel m.p.
1400
1200
T
max
°C
1300
1100
offset=1mm
offset=0
1000
900
0
10
20
30
40
50
X (mm)
FIGURE 5.5: Calculated maximum temperatures at the Ni-invar interface.
Values from points directly below the laser path and at 1mm offset.
It can be seen that the nodes directly under the laser path generally reach the
melting temperature, with some exceptions near the material interfaces as those
are reproduced by the model (Figure 5.1). Furthermore, the nodes that are
slightly offset reach a maximum temperature below the melting point, especially
when they are located directly over the copper layer. For nodes close to the layer
interfaces the situation is even more acute. The most obvious explanation lies
with the fact that some of these nodes lie at a larger depth from the surface than
those, for example, in the central 20-30mm region.
A concern for the actual embedding process is the maximum temperature that is
reached by the stainless steel substrate during laser deposition. The reason for
112
CHAPTER 5: MODELING
this concern is dual. First, the thin-film structure lies at the copper-steel interface
and it should not be exposed at high temperatures – so that thermal mismatch
strains are minimized – or steep temperature gradients – so that thermal shock is
avoided. Second, the stainless steel should not be treated for long times at
elevated temperatures (>700°C), so that precipitation of chromium at the grain
boundaries, and the subsequent loss of corrosion resistance are avoided
900
offset=0
Temperature (°C)
800
offset=2mm
700
600
500
400
300
1
1.5
2
2.5
3
3.5
4
4.5
5
Time (s)
FIGURE 5.6: Calculated temperature at the steel-copper interface.
Values from a point which lies directly below the laser path.
The calculated data shown in Figure 5.6, imply that the heating rate reaches a
maximum of 100°Cs-1, and the 700°C threshold is crossed for a very short time
(in the order of ~0.5s) and for the region directly below the laser beam only.
The experimental results have proven that the thin-film structure and the
enclosed sensor survive the heat exposure. If the results in Figure 5.4 can provide
any confidence to the validity of the simulation, then it can be inferred that the
predicted temperatures and heating rates are indeed close to reality.
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CHAPTER 5: MODELING
5.3 Summary
This chapter included the description and the results of a finite element analysis
model which was used to calculate temperature profiles at points of interest in
the multilayer structure containing the thin-film sensors. It was attempted to
create a model that would represent the geometry of the actual substrate prior
and during the embedding process and that would also simulate the moving
heat source of the Nd:YAG laser used in the processing. The mesh density was
adjusted so as to minimize non-physical temperature fluctuations and allow the
completion of one laser pass at a smaller-than-100h time-frame.
Simplicity of calculations dictated the use of temperature-independent material
properties, and the exclusion of convective and radiation cooling effects. The
boundary conditions imposed were defined by the heat input of the laser on part
of the surface, the contact with a heat sink at the bottom of the substrate, and the
contact with a volume of powder (itself in contact with a heat sink) at the sides of
the substrate. All other surfaces were considered insulated.
The temperature profile of a point at the nickel-invar interface was calculated
and the results were compared with experimental data. The qualitative
agreement was satisfactory, and the discrepancies between calculation and
experiment could be accounted for by the simplifying assumptions of the model.
The calculations also predicted that there is remelting at the invar-nickel interface
(as observed experimentally), and that the maximum temperature at the coppersteel interface does not exceed 700°C for more than 0.5s. These results provide
confidence that the model can offer insight for the effect of the protective layers’
thickness on the maximum temperature at the copper-steel interface.
Furthermore, a more refined model in terms of material properties and mesh
density can be used to provide more accurate data. Use of this model’s results to
perform a stress analysis on the structure based on thermal expansion coefficients
114
CHAPTER 5: MODELING
and elastic moduli is currently limited by the time and space necessary to
complete the calculations.
115
6
Conclusions
The research described in the previous chapters focuses on the development of a
technology that allows the inclusion of thermo-mechanical sensors in the body of
metallic structures (tools, components, mechanisms). The inspiration for this
work came from the need to monitor the temperature and deformation of
structures in real time, in demanding environments (manufacturing, drilling,
combustion, etc.)
As a response to the push for automation in manufacturing, the main processes
used were borrowed from fields with significant progress towards this goal. In
particular, the sensors and part of the necessary enclosure were deposited with
thin-film techniques, already widely used in the IC and coatings industries. The
structures were built using Shape Deposition Manufacturing (SDM) in
conjunction with Laser Deposition. SDM is a methodology that automates the
design and manufacturing of geometrically complex objects using additive and
subtractive techniques, whereas Laser Deposition is a process used to
incrementally fabricate metal objects with a high quality of properties.
Even though attempts to monitor temperature and strain with thin film sensors
were carried out in the past, mainly in turbomachinery facilities, certain things
were achieved for the first time, at the proof-of-concept level, during this
research.
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CHAPTER 6: CONCLUSIONS
Namely, it became possible to deposit pinhole-free dielectric layers on nonspecial commonly available substrates (e.g.: 304L stainless steel) with minimal
preparation, relative to the prior practice, and, most importantly, outside a clean
room environment. This accomplishment is generating hope that thin-film
deposition and shaping techniques could be integrated in a traditional
manufacturing environment without the elevated costs of maintaining particlefree areas.
The deposited strain sensors were tested in a four-point bend configuration and
their response was perfectly linear in the elastic regime. The output of the
experimental strain gage was also compared to that of a commercially available,
calibrated sensor, which was loaded under the same conditions. The comparison
provided a calibration for the deposited sensor and verified its linear response,
and, most importantly, the absence of hysteresis – a crucial feature for a strain
sensor.
In addition, the dielectric layers were grown directly on the substrates, without
treating coatings that form thermally grown oxides, and with deposition rates
significantly higher (~10 times) than those achieved by more expensive
techniques (i.e.: rf sputtering). In particular, all aluminum oxide films were
deposited using pulsed magnetron reactive sputtering, with maximum rates
close to 0.8nms-1 at a power density of ~12Wcm-2. The films were amorphous,
since they were deposited on cool substrates (~100°C), yet their dielectric
strength was within the range of reported values for aluminum oxide.
Most importantly, the sensors, after being enclosed in an insulating envelope,
were embedded in a metallic structure using high-temperature techniques (i.e.:
Laser Deposition). Since the particular technique creates a molten pool on a metal
powder bed it was essential to protect the sensitive thin-film structure during the
process. This was achieved with a combination of protective layers which were
117
electroplated on the substrate and partially buried the sensor and its insulation.
The protective layers were composed of copper and nickel.
Finally, a finite element model was created in order to simulate the embedding
process and thereby to provide insight and predictive capability with respect to
the geometry of the protective layers (size and thickness) and the parameters of
the laser (speed, pattern, power). The results of the model, which were compared
with experimental temperature measurements and structure observations,
proved encouraging for the applicability of the simulation not only for the
embedding of sensors, but for the general additive process as well.
At the current state there are a number of obstacles that still have to be overcome
so that a fully functional tool or component can be manufactured and put in
service. One of the most apparent is the shaping of the thin films over curved
surfaces, such as those produced by 5-axis machining. This feature will be vital in
order to locate the contact pads for the sensors in a position far from the area that
will engage in a harsh environment.
Another very important issue is that of splicing and telemetry. For most cases it
will be necessary to connect the sensor to an external signal processing device.
This will have to happen at the contact pads, therefore the connections must be
strong, reliable, and stable under high temperatures and probably embedded in a
protective layer themselves. Advances in telemetry will allow the transfer of the
signal to remote locations without the need for extensive wiring of some sort. In
the extreme case where a rough signal processing circuit is fabricated and
embedded inside the tooling, the results of such processing will still have to be
communicated to remote systems.
On another front, the act of embedding a structure inside another one
exacerbates the issues already faced by many layered manufacturing techniques.
Not only is there the problem of residual thermal stresses which affect the
integrity of the structure, there is an extra discontinuity as well. This
118
discontinuity might act as a stress concentration, thus affecting the fatigue limit
of the component or tool. Since “inserted“ structure will generally be very thin, it
will be a favorable site for crack initiation in an already non-continuous
structure.
Some of the problems outlined above can be remedied by using other type of
sensors in order to measure thermo-mechanical response. Currently, at the Rapid
Prototyping Laboratory there is active research in the area of embedding opticalfiber sensors in layered manufactured tools. Such sensors have the distinctive
advantage of non-contact signal transfer as can be accessed by optical means.
More importantly, they have no need for electrical insulation as their mode of
operation is based on optical properties, and therefore they are more robust in
terms of functionality. However, their placement cannot be easily automated or
parallelized.
119
A
The thin-film deposition system
Our thin film deposition system consists of a stainless steel cylindrical chamber
of 62 l capacity (diam:46cm, height:38cm), with two K.J. Lesker 1-KW magnetron
guns. The chamber is evacuated with a Leybold TMP361 turbomolecular pump
with a nominal maximum pumping speed of 400 l/s (for nitrogen) and capable
of achieving an ultimate pressure of less than 1E-8 Pa. The turbomolecular pump
is backed by an Alcatel Pascal 2021 mechanical pump with a pumping speed of
69 l/s and capable of achieving an ultimate pressure of less than 0.2 Pa. The
mechanical pump is also used to pump the chamber down to a rough vacuum of
30 Pa before the gate valve to the turbomolecular pump is opened.
It is suggested that the pressure at the foreline of the turbo pump remain in the
0.1-1 Pa range and never exceed 50 Pa. Moreover, it is strongly advised to avoid a
very low pressure either at the foreline or the gate to the mechanical pump since
this might lead to oil backstreaming from the mechanical pump. Oil
backstreaming becomes a severe problem in the transition flow regime which is
characterized by Knudsen numbers in the 0.01-1 range. The Knudsen number
Kn=l/L, is the ratio of the mean free path l of a gas molecule to a characteristic
distance L of the vacuum system. Therefore, for the foreline and roughing lines
of our system, with a diameter of 3 cm, the mean free path should always be less
than 0.03 cm to ensure that the lines are in the fluid flow regime. The mean free
path (in cm) for room temperatures is roughly inversely proportional to the gas
120
APPENDIX A
pressure (in Pa) [D. Smith], hence a pressure of at least 33 Pa is required at those
lines before they are valved off. This is easy to achieve for the roughing line to
the main chamber but difficult to maintain at the foreline during long
pumpdown procedures. In such cases the injection (purge) of argon or nitrogen
at the foreline is suggested so that the pressure stays within desired limits. We
have not implemented such measures as of this writing.
A recirculating water chiller is used to provide cool water (21 ˚C) to the
magnetron guns and the turbo pump. The guns should operate at a maximum
temperature of 25 ˚C and should be cooled with water flowing at 1.9 l/min. The
pump should operate in the 10-30 ˚C range and the minimum suggested flow is
0.3 l/m. Our Tek-Temp TKD-100 chiller is capable of pumping 6.5 l/min at 550
KPa with a cooling capacity of 2 KW. Currently it operates at 300-350 KPa.
The magnetron guns face upwards and accept disk-shaped targets which are 3
mm thick and 50 mm in diameter. The targets are clamped mechanically on the
copper backing plate which is cooled by recirculating water and the plasma is
confined on top of them with the aid of titanium ground shields. The guns are
driven by an ENI RPG-50 5 KW power supply which can deliver a dc or pulsed
dc signal at a preset level of power, voltage or current. The frequency of the
pulsed dc signal can be varied between 50 and 250 KHz with a pulse width in the
400-3000 ns range, depending on the operating frequency. The principle of
operation for the pulsed dc mode is outlined in a separate appendix [or
chapter?].
The working gases are introduced into the chamber via a mass flow control
system by MKS. The flow rate for argon and oxygen is adjusted by two mass
flow controllers that have been calibrated for the particular gases. The pressure at
the controller inlet is maintained at approximately 700 KPa with a gas regulator.
The working pressure is measured with a MKS 127A Baratron capacitance
manometer with a dynamic range of 1E5 and an accuracy of 1.33E-4 Pa. It is
121
APPENDIX A
suggested that the Baratron sensor is valved off during venting to avoid
exposure to higher pressures and that its zero offset is corrected before every
deposition. Before such a correction it is suggested that the Baratron remain at a
pressure that is 2 orders of magnitude below its precision limit for 4-6 hours. The
zero offset can be adjusted either mechanically with a set screw at the back of the
sensor or digitally by the mass flow control system. The base pressure is
measured by a Bayard-Alpert ionization gage (G100F by K.J. Lesker). This gage
has an upper working pressure limit of 0.13 Pa and its emission current must
always be switched off when pressure gets near this value. It is advisable that the
gage is degassed for a few minutes before measuring the base pressure. The
degassing procedure is most effective when it is performed during a bakeout and
can only be initiated if the pressure is already below 1.3E-3 Pa. High pressures,
mainly during venting and roughing procedures, are measured by a
thermocouple pressure sensor. Such a sensor is also used to monitor the pressure
at the foreline of the turbomolecular pump. Those sensors have a sensitivity low
limit of 0.13 Pa.
The chamber can be “baked” to accelerate the desorption of water molecules
from the internal surfaces with the aid of 627 W heating tape that is wrapped
around it. The chamber walls are also wrapped with aluminum foil to reduce the
heat being lost by re-radiation to the environment. The heating tape is attached to
a temperature control system but even when it is driven at full capacity, the
internal surface temperature does not exceed 85 ˚C. With a 10-hour bake-out
followed by cooling for another 10-12 hours the turbomolecular pump can
achieve pressures as low as 4E-5 Pa. This is quite good for a system that has 20
ports sealed with small O-rings and also a large O-ring that seals the top plate
itself. Out of the 13 through-holes on the baseplate, 2 are used for the power
feedthroughs to the magnetron guns, 1 for a type-K thermocouple feedthrough, 1
for a power feedthrough to the substrate holder, and 1 for a rotational
feedthrough to the mechanical shutter. There is also one O-ringed seal for the
push-pull transport feedthrough for the substrate holder.
122
APPENDIX A
The substrate holder is a construction of two parallel stainless steel plates which
are connected with teflon spacers and nylon screws to ensure electrical isolation.
The bottom plate has a number of threaded holes so that substrates can be
fastened to it. It is placed on an aluminum rack system, isolated from it by teflon
sliders, and can be translated from one gun to another with the aid of the
aforementioned push-pull transport mechanism. The latter is attached to the top
plate which serves also as a ground shield during sputter-etching of the
substrates. The substrate holder is driven by the ENI power supply via a
shielded copper wire during sputter-etching and is at a floating potential during
deposition.
123
B
Positive masks of strain sensors
50
75
125
100
420
330
FIGURE B.1: The numbers denote the smallest linewidth in µm.
124
C
Sample LabView file
The file below controls the laser and the robotic arm that carries it during the
embedding process:
1
13 2 0 135 0.00 0.00 0.00
14 0.00 0.00 0.00 0.0 0.0 0.0
12 0 -35 0.0 0.0 20
11 10000
19 8
17 0 0 0
18 1 0 0
21 3
20 1
2 0 29 0.0 0.0 15
2 0 29 0.0 0.0 15
20 0
12 0 29 0.0 0.0 15
11 20000
17 0 0 0
12 17.9 35 0.0 0.0 15
11 3000
20 1
2 17.9 -36 0.0 0.0 15
2 17.9 -36 0.0 0.0 15
17 0 0 0
20 0
12 17.9 -36 0.0 0.0 15
11 20000
12 1.8 -24 0.0 0.0 15
11 1000
20 1
2 1.8 24 0.0 0.0 15
2 1.8 24 0.0 0.0 15
20 0
;
;
;
;
;
;
;
;
;
set's part center relative to pallet
tool offset
send and execute move
wait on Labview side
set laser power (10=full power)
set current powder feeder rates (10=max.)
turn on ALL powder feeders to be used
wait on ADEPT side (during moves)
open shutter
; PASS 1
; PASS 2
; PASS 3
125
APPENDIX C
12 1.8 24 0.0 0.0 15
11 20000
12 3.6 24 0.0 0.0 15
11 1000
20 1
2 3.6 -24 0.0 0.0 15
2 3.6 -24 0.0 0.0 15
20 0
12 3.6 -24 0.0 0.0 15
11 20000
12 5.4 -24 0.0 0.0 50
11 1000
20 1
2 5.4 24 0.0 0.0 15
2 5.4 24 0.0 0.0 15
20 0
12 5.4 24 0.0 0.0 15
11 20000
12 7.2 24 0.0 0.0 15
11 1000
20 1
2 7.2 -24 0.0 0.0 15
2 7.2 -24 0.0 0.0 15
20 0
12 7.2 -24 0.0 0.0 15
11 20000
12 9 -24 0.0 0.0 50
11 1000
20 1
2 9 24 0.0 0.0 15
2 9 24 0.0 0.0 15
20 0
12 9 24 0.0 0.0 15
11 20000
12 10.8 24 0.0 0.0 15
11 1000
20 1
2 10.8 -24 0.0 0.0 20
2 10.8 -24 0.0 0.0 20
20 0
12 10.8 -24 0.0 0.0 20
11 20000
12 12.6 -24 0.0 0.0 20
11 1000
20 1
2 12.6 24 0.0 0.0 20
2 12.6 24 0.0 0.0 20
20 0
12 12.6 24 0.0 0.0 20
11 20000
12 14.4 24 0.0 0.0 20
11 1000
20 1
; PASS 4
; PASS 5
; PASS 6
; PASS 7
; PASS 8
; PASS 9
126
APPENDIX C
2 14.4 -24 0.0 0.0 25
2 16.2 -24 0.0 0.0 25
2 16.2 24 0.0 0.0 25
2 16.2 24 0.0 0.0 25
19 0
20 0
12 16.2 25 0 0 25
11 2000
0
; PASS 10
; PASS 11
127
D
ABAQUS heat transfer input file
This file is used to build the model described in “Modeling” on page 100 and
simulate the embedding process.
*HEADING
Heat transfer analysis for sample with 304L base (50x50x3)
copper, nickel and invar layers
TRANSIENT HEAT TRANSFER
UNITS: SI (m,kg,s,K)
**
**
**
APEX DEFINITION FOR SOLID MODEL (7 mm HIGH)
**
*NODE
999,0,0,0
1043,.055,0,0
3499,0,.050,0
3543,.055,.050,0
42999,0,0,.007
43043,.055,0,.007
45499,0,.050,.007
45543,.055,.050,.007
**
**
**
**
NODE GENERATION FOR 304L + COPPER + NICKEL + INVAR
**
**
44 INTERVALS ALONG X
**
**
X-EDGE AT 0,0-50,0 (Z=0)
**
*NGEN, NSET=XEDGE1
999,1043,1
**
**
X-EDGE AT 0,50-50,50 (Z=0)
128
APPENDIX D
**
*NGEN, NSET=XEDGE2
3499,3543,1
**
**
X-EDGE AT 0,0-50,0 (Z=7)
**
*NGEN, NSET=XEDGE3
42999,43043,1
**
**
X-EDGE AT 0,50-50,50 (Z=7)
**
*NGEN, NSET=XEDGE4
45499,45543,1
**
**
**
50 INTERVALS ALONG Y
**
**
BASE (Z=0)
**
*NFILL, NSET=BASE
XEDGE1,XEDGE2,50,50
**
**
TOP
**
*NFILL, NSET=TOP (Z=7)
XEDGE3,XEDGE4,50,50
**
**
**
BODY OF SUBSTRATE+LAYERS
**
*NFILL, NSET=NALL
BASE,TOP,14,3000
**
**
**
**
**
BOUNDARY DEFINITION
*NSET,NSET=OUTBOUND,GENERATE
3999,6499,50
6999,9499,50
9999,12499,50
12999,15499,50
15999,18499,50
18999,21499,50
21999,24499,50
24999,27499,50
27999,30499,50
30999,33499,50
33999,36499,50
36999,39499,50
6499,6543
9499,9543
12499,12543
129
APPENDIX D
15499,15543
18499,18543
21499,21543
24499,24543
27499,27543
30499,30543
33499,33543
36499,36543
39499,39543
4043,6543,50
7043,9543,50
10043,12543,50
13043,15543,50
16043,18543,50
19043,21543,50
22043,24543,50
25043,27543,50
28043,30543,50
31043,33543,50
34043,36543,50
37043,39543,50
**
**
**
**
**
ELEMENT DEFINITION FOR 304L
**
*ELEMENT, TYPE=DC3D8
1001,1001,1002,1052,1051,4001,4002,4052,4051
**
**
ELEMENT GENERATION FOR 304L (6 LAYERS, EACH 500 MICRONS THICK)
**
*ELGEN, ELSET=ESTEEL
1001,40,1,1,50,50,50,6,3000,3000
**
**
**
ELEMENT DEFINTIION FOR 2ND LAYER
**
20001,20001,20002,20052,20051,23001,23002,23052,23051
**
**
**
ELEMENT GENERATION FOR 2nd LAYER
**
*ELGEN
20001,40,1,1,30,50,50,2,3000,3000
**
**
**
**
ELEMENT SET GENERATION FOR COPPER IN 2ND LAYER
**
*ELSET, ELSET=ECOPPER, GENERATE
20013,20028
130
APPENDIX D
20063,20078
20113,20128
20163,20178
20213,20228
20263,20278
20313,20328
20363,20378
20413,20428
20463,20478
20513,20528
20563,20578
20613,20628
20663,20678
20713,20728
20763,20778
20813,20828
20863,20878
20913,20928
20963,20978
23013,23028
23063,23078
23113,23128
23163,23178
23213,23228
23263,23278
23313,23328
23363,23378
23413,23428
23463,23478
23513,23528
23563,23578
23613,23628
23663,23678
23713,23728
23763,23778
23813,23828
23863,23878
23913,23928
23963,23978
**
**
**
**
ELEMENT SET GENERATION FOR NICKEL IN 2ND LAYER
**
*ELSET, ELSET=ENICKEL2, GENERATE
20005,21455,50
20006,21456,50
20007,21457,50
20008,21458,50
20009,21459,50
20010,21460,50
20011,21461,50
20012,21462,50
131
APPENDIX D
20029,21479,50
20030,21480,50
20031,21481,50
20032,21482,50
20033,21483,50
20034,21484,50
20035,21485,50
20036,21486,50
21013,21028
21063,21078
21113,21128
21163,21178
21213,21228
21263,21278
21313,21328
21363,21378
21413,21428
21463,21478
23005,24455,50
23006,24456,50
23007,24457,50
23008,24458,50
23009,24459,50
23010,24460,50
23011,24461,50
23012,24462,50
23029,24479,50
23030,24480,50
23031,24481,50
23032,24482,50
23033,24483,50
23034,24484,50
23035,24485,50
23036,24486,50
24013,24028
24063,24078
24113,24128
24163,24178
24213,24228
24263,24278
24313,24328
24363,24378
24413,24428
24463,24478
**
**
**
**
ELEMENT SET GENERATION FOR INVAR IN 2ND LAYER
**
*ELSET, ELSET=EINVAR2, GENERATE
20001,21451,50
20002,21452,50
20003,21453,50
132
APPENDIX D
20004,21454,50
20037,21487,50
20038,21488,50
20039,21489,50
20040,21490,50
23001,24451,50
23002,24452,50
23003,24453,50
23004,24454,50
23037,24487,50
23038,24488,50
23039,24489,50
23040,24490,50
**
**
**
**
**
ELEMENT DEFINITION FOR 3RD LAYER
**
26003,26003,26004,26054,26053,29003,29004,29054,29053
**
**
**
ELEMENT GENERATION FOR 3rd LAYER
**
*ELGEN
26003,36,1,1,30,50,50,2,3000,3000
**
**
**
**
**
ELEMENT SET GENERATION FOR NICKEL IN 3RD LAYER
**
*ELSET, ELSET=ENICKEL3, GENERATE
26011,26030
26061,26080
26111,26130
26161,26180
26211,26230
26261,26280
26311,26330
26361,26380
26411,26430
26461,26480
26511,26530
26561,26580
26611,26630
26661,26680
26711,26730
26761,26780
26811,26830
26861,26880
26911,26930
133
APPENDIX D
26961,26980
27011,27030
27061,27080
27111,27130
27161,27180
27211,27230
27261,27280
27311,27330
27361,27380
27411,27430
27461,27480
29011,29030
29061,29080
29111,29130
29161,29180
29211,29230
29261,29280
29311,29330
29361,29380
29411,29430
29461,29480
29511,29530
29561,29580
29611,29630
29661,29680
29711,29730
29761,29780
29811,29830
29861,29880
29911,29930
29961,29980
30011,30030
30061,30080
30111,30130
30161,30180
30211,30230
30261,30280
30311,30330
30361,30380
30411,30430
30461,30480
**
**
**
**
ELEMENT SET GENERATION FOR INVAR IN 3RD LAYER
**
*ELSET,ELSET=EINVAR3, GENERATE
26003,27453,50
26004,27454,50
26005,27455,50
26006,27456,50
26007,27457,50
26008,27458,50
134
APPENDIX D
26009,27459,50
26010,27460,50
26031,27481,50
26032,27482,50
26033,27483,50
26034,27484,50
26035,27485,50
26036,27486,50
26037,27487,50
26038,27488,50
29003,30453,50
29004,30454,50
29005,30455,50
29006,30456,50
29007,30457,50
29008,30458,50
29009,30459,50
29010,30460,50
29031,30481,50
29032,30482,50
29033,30483,50
29034,30484,50
29035,30485,50
29036,30486,50
29037,30487,50
29038,30488,50
**
**
**
**
ELEMENT DEFINITION FOR 4TH LAYER (INVAR ONLY)
**
38009,32009,32010,32060,32059,35009,35010,35060,35059
**
**
**
ELEMENT GENERATION FOR INVAR IN 4TH LAYER (2 layers 500 microns each)
**
*ELGEN, ELSET=EINVAR4
38009,24,1,1,30,50,50,2,3000,3000
**
**
**
ELEMENT DEFINITION FOR OUTER-LEFT LAYER (POWDER ONLY)
**
999,999,1000,1050,1049,3999,4000,4050,4049
**
**
**
ELEMENT GENERATION FOR POWDER IN 4TH LAYER (2 layers 500 microns each)
**
*ELGEN, ELSET=OUT1
999,2,1,1,50,50,50,12,3000,3000
**
**
135
APPENDIX D
**
ELEMENT DEFINITION FOR OUTER-RIGHT LAYER (POWDER ONLY)
**
1041,1041,1042,1092,1091,4041,4042,4092,4091
**
**
**
ELEMENT GENERATION FOR POWDER IN 4TH LAYER (2 layers 500 microns each)
**
*ELGEN, ELSET=OUT2
1041,2,1,1,50,50,50,12,3000,3000
**
**
**
**
ELEMENT SET GENERATION FOR NICKEL AND POWDER
**
*ELSET, ELSET=ENICKEL
ENICKEL2,ENICKEL3
**
*ELSET, ELSET=EINVAR
EINVAR2,EINVAR3,EINVAR4
**
**
*ELSET, ELSET=EPOWDER
OUT1,OUT2
**
**
**
**
MATERIAL PROPERTIES ETC.
**
**
**
*PHYSICAL CONSTANTS, ABSOLUTE ZERO=-273.15
**
**
*SOLID SECTION, MATERIAL=304L,ELSET=ESTEEL
*MATERIAL, NAME=304L
*CONDUCTIVITY
14.9
*DENSITY
7900
*SPECIFIC HEAT
477
**
**
*SOLID SECTION, MATERIAL=COPPER, ELSET=ECOPPER
*MATERIAL, NAME=COPPER
*CONDUCTIVITY
401
*DENSITY
8933
*SPECIFIC HEAT
385
**
136
APPENDIX D
**
*SOLID SECTION, MATERIAL=NICKEL, ELSET=ENICKEL
*MATERIAL, NAME=NICKEL
*CONDUCTIVITY
90.7
*DENSITY
8900
*SPECIFIC HEAT
444
**
**
**
**
INVAR (SOLID)
**
*SOLID SECTION, MATERIAL=INVARSOL, ELSET=EINVAR
*MATERIAL, NAME=INVARSOL
**
**
*CONDUCTIVITY
11
** DENSITY DEFINED AS 100% OF THEORETICAL
*DENSITY
8000
** SPECIFIC HEAT FROM
** ASM METALS HANDBOOK, Low-Expansion alloys
*SPECIFIC HEAT
515
** LATENT HEAT IS THE ARITHMETIC MEAN FOR 64FE-36NI
*LATENT HEAT
282290,1440,1450
**
**
**
**
INVAR (POWDER)
**
*SOLID SECTION, MATERIAL=INVARPDR, ELSET=EPOWDER
*MATERIAL, NAME=INVARPDR
**
**
*CONDUCTIVITY
11
** DENSITY DEFINED AS 50% OF THEORETICAL
*DENSITY
4000
*SPECIFIC HEAT
515
**
**
**
**
INITIAL CONDITIONS
*INITIAL CONDITION, TYPE=TEMPERATURE
NALL,320
**
137
APPENDIX D
**
**
AMPLITUDE DEFINITIONS
**
**
LASER SPEED = 15 mm/sec
**
*AMPLITUDE, NAME=E1, TIME=STEP TIME
0,0,0.167,1,0.333,0
0.167,0,0.333,1,0.5,0
0.333,0,0.5,1,0.667,0
0.5,0,0.667,1,0.834,0
0.667,0,0.834,1,1.,0
0.834,0,1.,1,1.167,0
1.,0,1.167,1,1.333,0
1.167,0,1.333,1,1.5,0
1.333,0,1.5,1,1.667,0
1.5,0,1.667,1,1.834,0
1.667,0,1.834,1,2.,0
1.834,0,2.,1,2.167,0
2.,0,2.167,1,2.333,0
2.167,0,2.333,1,2.5,0
2.333,0,2.5,1,2.667,0
2.5,0,2.667,1,2.834,0
2.667,0,2.834,1,3.,0
2.834,0,3.,1,3.167,0
3.,0,3.167,1,3.333,0
3.167,0,3.333,1,3.5,0
**
**
**
**
DEFINITION OF ELEMENT SETS
**
FOUR ELEMENTS MAKE UP AN ELEMENT SET WITH SIZE 2X2.5 mm
**
**
*ELSET,ELSET=S101421
138
APPENDIX D
23451,23452
23401,23402
29453,29454
29403,29404
29455,29456
29405,29406
29457,29458
29407,29408
41459,41460
41409,41410
41461,41462
41411,41412
41463,41464
41413,41414
41465,41466
41415,41416
41467,41468
41417,41418
41469,41470
41419,41420
41471,41472
41421,41422
41473,41474
41423,41424
41475,41476
41425,41426
41477,41478
41427,41428
41479,41480
41429,41430
41481,41482
41431,41432
29483,29484
29433,29434
29485,29486
29435,29436
139
APPENDIX D
29487,29488
29437,29438
23489,23490
23439,23440
**
**
**
**
**
OUTPUT DEFINITIONS
**
*PREPRINT, MODEL=NO,HISTORY=NO
*RESTART,WRITE,OVERLAY
**
**
**
**
STEP DEFINITIONS
**
**
**
STEP 2
**
*STEP, INC=2000
*HEAT TRANSFER,END=PERIOD,DELTMX=1500
.002,20,.00001
**
**
**
BOUNDARY CONDITION FOR BASE (T=350 CELSIUS)
**
BOUNDARY CONDITION FOR OUTER INVAR POWDER (T=320 CELSIUS)
**
*BOUNDARY
BASE,11,,350
OUTBOUND,11,,320
**
**
**
**
*DFLUX,AMPLITUDE=E1
S101440,S2,3.3E8
*DFLUX,AMPLITUDE=E2
S101439,S2,3.3E8
*DFLUX,AMPLITUDE=E3
S101438,S2,3.3E8
*DFLUX,AMPLITUDE=E4
S101437,S2,3.3E8
*DFLUX,AMPLITUDE=E5
S101436,S2,3.3E8
*DFLUX,AMPLITUDE=E6
S101435,S2,3.3E8
*DFLUX,AMPLITUDE=E7
S101434,S2,3.3E8
*DFLUX,AMPLITUDE=E8
S101433,S2,3.3E8
140
APPENDIX D
*DFLUX,AMPLITUDE=E9
S101432,S2,3.3E8
*DFLUX,AMPLITUDE=E10
S101431,S2,3.3E8
S101430,S2,3.3E8
S101429,S2,3.3E8
S101428,S2,3.3E8
S101427,S2,3.3E8
S101426,S2,3.3E8
S101425,S2,3.3E8
S101424,S2,3.3E8
S101423,S2,3.3E8
S101422,S2,3.3E8
S101421,S2,3.3E8
**
**
**
*NODE FILE
NT
*END STEP
141
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145

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