Design, Fabrication, and Characterization of a MEMS
Steam-Generating Device Based on the Decomposition of
High-Test Hydrogen Peroxide
by
FERAS EID
Master of Science in Mechanical Engineering
Massachusetts Institute of Technology, 2006
]
MASSACHUSETTS INS IUTE
OFTECHNOLOG
I
SEP 0 12010
LIBRARIE,
Bachelor of Engineering in Mechanical Engineering
American University of Beirut, 2004
ARCHIVES
Submitted to the Department of Mechanical Engineering
in Partial Fulfillment of the Requirements for the Degree of
DOCTOR OF PHILOSOPHY IN MECHANICAL ENGINEERING
at the
Massachusetts Institute of Technology
June 2010
© 2010 Massachusetts Institute of Technology
All rights reserved
A
Sign atu re of Au th or ............................................................................................................
........................
Department of Mechanical Engineering
May 19, 2010
Certified by ....................................................................
Carol Livermore
Associate Professor of Mechanical Engineering
Thesis Supervisor
Accepted by ...............................................................
David Hardt
Chairman, Department Committee on Graduate Students
2
Design, Fabrication, and Characterization of a MEMS Steam-Generating
Device Based on the Decomposition of High-Test Hydrogen Peroxide
by
Feras Eid
Submitted to the Department of Mechanical Engineering on May 19, 2010
in Partial Fulfillment of the Requirements for the Degree of
Doctor of Philosophy in Mechanical Engineering
Abstract
Microscale ejector pumps offer the potential for high flow rate pumping of gases, a
functionality that is greatly needed in MEMS technology. These pumps have many
additional characteristics, such as their simplicity of design and their lack of moving parts,
which favor them over other state-of-the-art MEMS gas pumps. One of the challenges
associated with driving ejector pumps, however, is providing a compact source of motive
fluid. This fluid is the high-speed gas that drives the pumping action. The current thesis
presents a MEMS device capable of generating steam at speeds suitable for driving an
ejector pump in a compact fashion. To that end, the device utilizes the homogeneous
catalytic decomposition of hydrogen peroxide. Analysis shows that a MEMS ejector pump
driven by this device is capable of handling mass flow rates per unit pump volume on the
order of 10-2 g/s/cm 3, which are two orders of magnitude higher than those of state-of-theart MEMS gas pumps. In addition to pumping, the steam generator may also be used for
microrocket thrust generation in micropropulsion applications.
In this thesis, the design, fabrication, testing, and successful demonstration of the
MEMS steam generator are presented. The device consists of a mixing section for the
peroxide and catalyst streams, a reactor section where the peroxide decomposes, and
finally a nozzle section where the gaseous products of the decomposition are accelerated to
the required velocities. To design the device, multidomain (chemical, thermal, and fluidic)
numerically-implemented modeling is used to study the underlying physics and arrive at
an optimized, microfabricatable design. The modeling takes into account the key challenges
of thermal management, achieving fast mixing, and boundary layer compensation. The
device is then fabricated from a stack of four silicon wafers and one Pyrex wafer using deep
reactive ion etching and wafer bonding. The modeling also guides the design of a micabased ceramic package which provides both thermal insulation and piping ports. The
system is then experimentally tested using high-test hydrogen peroxide and ferrous
chloride tetrahydrate solution as the catalyst. The overall initial peroxide mass fraction is
varied between 83% and 71%. The device is characterized using temperature
measurements, refractive index analysis, and visual inspection during operation.
Successful performance is demonstrated via the full decomposition of the peroxide and the
complete vaporization of the water produced. The experimental results are also compared
with those from the simulation. Good agreement is observed between experiment and
theory, providing comprehensive model verification. The realization and demonstration of
this steam generator promise significant enhancements in MEMS technology, particularly
in the fields of gas pumping and micropropulsion.
Thesis Supervisor: Carol Livermore
Title: Associate Professor, Mechanical Engineering
Acknowledgements
I feel very lucky for having had such an enjoyable and rewarding PhD experience. I
owe a lot of that to the support of my family and friends, but also to the good fortune of
being part of an amazing research group ever since I joined MIT. My supervisor, Prof. Carol
Livermore, has been a great mentor and friend. I cannot thank her enough for her
continuous support and guidance, for always being available to discuss any questions or
concerns I had, and for her constant cheerfulness and understanding, which can make a
world of difference in a student's graduate school experience. My labmates have also been
an integral part of this experience. Thank you Frances, Gunjan, Nader, Lei, Aalap, and Eric
for all the fun outings, get-togethers, and surprise birthday parties. I will definitely miss
those times! Thanks in particular to my officemates Frances, Gunjan, and Lei for all the
conversations and laughs, and for their willingness to listen to my rants when the research
was facing some obstacles.
I also thank Dr. Luis Fernando Veldsquez-Garcia for the immense technical help he
provided during the different stages of the project. One aspect of this work that I
particularly pride myself on is that the very first generation of devices that I fabricated
worked as expected. This is largely due to Luis's helpful suggestions and willingness to
answer my numerous questions, especially during the fabrication stage. I also thank my
committee members, Prof. Jeffrey Lang and Prof. Evelyn Wang, for their helpful ideas
during our meetings. My thanks also extend to our administrative assistant Natalie Weaver
for her help in ordering parts that I needed for the project, to the MTL staff for fab training
and consultations, to Kimberlee Collins for her CFD simulation of the mixer, to Frances Hill
for her help in taking the SEM image, to Tyrone Hill for showing me the components of his
test setup, and to Daniel Herrick, Andrew Kalil, and Steven Yuan for their help in ensuring
safety during experiments. The project was financially supported by Prof. Livermore, the
Mechanical Engineering Department at MIT (through teaching assistantships), and partially
by DARPA, MDA, and AFRL.
I'd also like to acknowledge some friends outside MIT who have made my life during
this period much more enjoyable: Joe, Amal, Lama, and Mounir, who have become my
"home away from home," Siddarth for being a great roommate, friend, and movie
companion, Levi for memorable gatherings and Thanksgiving dinners, and Andrew for
many interesting conversations and for his help when I was changing apartments.
Finally, I'd like to acknowledge my parents, sister, two brothers, and my siblings'
families, for their continuous moral support during all this time that I've spent abroad. Our
phone conversations and my summer visits to Lebanon have been my constant source of
recharging for the demanding PhD work. I look forward to seeing you all again soon!
Table of Contents
1 INTRODUCTION
1.1 Overview
14
1.2 Literature Review of MEMS Pumps
15
1.3 Ejector Pumps
18
1.3.1
1.3.2
1.3.3
1.3.4
Overview .....................................................................................................................................................
Operating principle ......................................................................................................................................
M acroscale characteristics and applications..........................................................................................
M icroscale ejector pum ps ...........................................................................................................................
1.4 Steam Generation from the Decomposition of High-Test Hydrogen Peroxide
1.4.1 Overview and advantages ...........................................................................................................................
1.4.2 Other examples and applications of high-test hydrogen peroxide decomposition...............................
2
14
18
18
19
20
23
23
25
1.5 Thesis Roadmap
28
MODELING AND DESIGN
31
2.1 Background and Challenges
31
2.2 Conceptual Design
34
2.3 Mixer Design
36
2.3.1 M ixer overview ............................................................................................................................
.......... 36
2.3.2 M icrom ixing literature review .......................................................................................
............................ 37
2.3.3 M ixer modeling and design .....................................................................................................................
39
2.3.4 CFD sim ulation of the mixer ......................................................................................................................
42
2.4 Reactor
42
2.4.1 Reactor overview .........................................................................................................................
42
2.4.2 Reactor modeling and design ...............................................................
.................................................... 43
2.5 Nozzle
53
2.5.1 Nozzle overview ................
......................................
.......... 53
2.5.2 Nozzle modeling and design........................................................................................................................53
2.6 Parametric Study
59
2.7 Overall Design
65
2.8 Thermal Management
2.9 Analysis of an Ejector Pump Based on the Current Work.
3 FABRICATION
72
79
3.1 Overview
3.2 Wafer-Level Die Layout
3.3 Alignment Marks
3.4 Layer 1
3.5 Layer 3
3.6 Layers 2 and 4
3.7 Wafer Bonding
4 TEST RIG SETUP
100
4.1 Overview
100
4.2 Hydrogen Peroxide and Safety
100
4.3 Package
103
4.4 Test Rig Components
108
4.5 Component Passivation_
112
5 TESTING
114
5.1 Overview
114
5.2 Experiments with Mixer-Testing Devices
114
5.3 Experiments with Nominally-Designed Devices
116
5.3.1 Experimental conditions..........................................................................................................................116
5.3.2 Visual inspection ..................................................................................................
117
5.3.3 Refractive index analysis ...........................................................................................................................
120
5.3.4 Device wall temperature measurements ..................................................................................................
123
5.3.5 Effluent temperature measurements .......................................................................................................
125
5.3.6 Summary of the implications of the experimental results........................................................................130
6 CONCLUSIONS
131
6.1 Summary and Important Findings
131
6.2 Challenges and Future Improvements
134
6.3 Design Modifications for Future Applications
138
6.4 Concluding Remarks
140
APPENDIX A: Detailed Process Flow
141
APPENDIX B: Practices Followed For Successful Wafer Bonding
145
REFERENCES
147
List of Figures
Figure 1.1. Schem atic of an ejector pum p ............................................................................
19
Figure 2.1. Conceptual design ...............................................................................................
35
Figure 2.2. Schematic of a mixer with dimensions .............................................................
40
Figure 2.3. Differential reactor elem ent ..............................................................................
48
Figure 2.4. Plot of the flow temperature along the reactor length for the design conditions,
with inserts magnifying the vaporization stages...................................................................
52
Figure 2.5. Plot of the species' mass fractions along the reactor length for the design
conditions, with an insert magnifying the region near the reactor inlet............................
52
Figure 2.6. Schematic of the nozzle showing the known quantities at different locations... 54
Figure 2.7. Differential nozzle elem ent..................................................................................
56
Figure 2.8. Nozzle width profile before and after boundary layer compensation with insert
show ing the throat region .....................................................................................................
58
Figure 2.9. Plot of static and stagnation pressures along the nozzle for the design conditions
.......................................................................................................................................................
59
Figure 2.10. Plot of the Mach number along the nozzle for the design conditions............ 59
Figure 2.11. Plot of peroxide mass fraction at the reactor exit versus initial peroxide mass
fra ctio n .........................................................................................................................................
60
Figure 2.12. Plot of static and stagnation temperatures at nozzle exit versus initial peroxide
m ass fractio n ................................................................................................................................
61
Figure 2.13. Energy balance on the entire reactor .............................................................
61
Figure 2.14. Plot of reactor pressure versus initial peroxide mass fraction ......................
64
Figure 2.15. 3D m odel of entire device.................................................................................
65
Figure 2.16. Schematic of a cross-section of the device at distance x along the flow direction,
illustrating the dominant heat transfer mechanisms for the configuration in which the
device is encased inside a package ........................................................................................
68
10
Figure 2.17. Thermal network between device and environment at steady state ............
68
Figure 2.18. Package design ..................................................................................................
71
Figure 2.19. O-ring gland design. Dimensions and tolerances are in mm. .........................
71
Figure 2.20. Schematic for ejector pump analysis................................................................
73
Figure 3.1. Schematic section-view of a microfabricated device ........................................
80
Figure 3.2. Process flow summary for transferring the alignment marks .........................
83
Figure 3.3. "Alignm ent Marks" m ask ...................................................................................
84
Figure 3.4. Rotated die-level zoom of the mask "Alignment Marks" ..................................
84
Figure 3.5. Zoorns of the wafer-bonding and the top-bottom alignment marks................ 85
Figure 3.6. Process flow sum mary for Layer 1 ....................................................................
86
Figure 3.7. "Holes" mask with insert showing the complementary top-bottom alignment
fe a ture s.........................................................................................................................................
87
Figure 3.8. Rotated die-level zoom of the mask "Holes"......................................................
88
Figure 3.9. Process flow summary for Layer 3 ....................................................................
89
Figure 3.10. "Deep Features" m ask......................................................................................
90
Figure 3.11. Rotated die-level zoom of the mask "Deep Features".....................................
90
Figure 3.12. Process flow summary for Layers 2 and 4 (continued on next page)............ 92
Figure 3.13. "All Features" m ask...........................................................................................
94
Figure 3.14. Rotated die-level zoom of the mask "All Features"......................... 94
Figure 3.15. Schematic of the fusion bonding process .........................................................
96
Figure 3.16. Schematic of the anodic bonding step .............................................................
97
Figure 3.17. Photograph of a microfabricated device with the nominal design................
98
Figure 3.18. Photograph of a microfabricated mixer-testing device having an extra outlet
port in the bottom for fluid discharge, which in the fully-functional devices occurs through
th e n ozzle .....................................................................................................................................
98
11
Figure 3.19. SEM image of the mixers in cross-sectional view, with an insert magnifying one
m ixer and show ing a wall protrusion....................................................................................
99
Figure 4.1. Peroxide storage in lab ..........................................................................................
102
Figure 4.2. Safety gear...............................................................................................................
102
Figure 4.3. Machined package made of Rescor 914 ...............................................................
105
Figure 4.4. Modified package bottom half for facilitating thermocouple insertion............. 106
Figure 4.5. Schematic explaining the difference in the thermocouple insertion methods
between the original and modified designs of the bottom half of the package ................... 107
Figure 4.6. Modified package bottom half for usage with the mixer-testing devices....... 108
Figure 4.7. Schem atic of test rig setup.....................................................................................
110
Figure 4.8. Assem bled package, fittings, and tubing ..............................................................
111
Figure 4.9. Photograph of the test rig highlighting the main components...........................
111
Figure 5.1. Frame-grabs of the device during experiment 1 showing the effluent in the
different stages: (a) startup, (b) intermediate transient period, and (c) steady state ....... 119
Figure 5.2. Procedure used to perform the refracive index analysis on the effluent ....... 122
Figure 5.3. Plot of the peroxide mass fraction at the reactor exit (from the simulation) and at
the nozzle exit (from experiments 1 and 2) versus initial peroxide mass fraction.............. 122
Figure 5.4. Plot of the silicon wall temperature from the simulation and the experiments
versus initial peroxide mass fraction ......................................................................................
124
Figure 6.1. Schematic of proposed modified setup with compressed air line for purging the
device after experim ents ..........................................................................................................
136
Figure 6.2. Photograph of the cracked bottom half of the package.......................................
137
Figure 6.3. Modified design for the bottom half of the package, which allows the use of a
m aterial with higher therm al conductivity.............................................................................
137
List of Tables
Table 1.1. Comparison of the specifications of some MEMS gas pumps in the literature with
those of an ejector pump based on the current work ........................................................
21
Table 2.1. Phases of H20 and H20 2 during the five reactor stages ......................................
44
Table 2.2. Variation of state variables during the different reactor stages........................
46
Table 2.3. Ejector pum p param eters .....................................................................................
78
Table 4.1. Comparison of the required package properties and those of Rescor 914 ........ 104
Table 5.1. Conditions during the peroxide experiments on nominally-designed devices.. 117
Table 5.2. Recovery factors, predicted adiabatic wall temperatures, and measured effluent
tem peratures for the three experim ents.................................................................................
129
CHAPTER ONE
1 INTRODUCTION
1.1 Overview
The past two decades have witnessed a marked increase in the interest in microelectro-mechanical systems (MEMS), both at the commercial and research levels.
Accelerometers, pressure transducers, inkjet printer cartridges, chemical and flow sensors,
and lab-on-chip devices for point-of-care medical testing are only a few examples of
systems that have been enabled or strongly improved by MEMS technology [1]. This
technology broadly refers to a wide range of fabrication methods that allow the massproduction of small-scale systems and components. Some of those fabrication methods
include photolithography, physical and chemical material deposition, wet and dry etching
for material removal, doping, chip and wafer bonding, and soft lithography (imprinting).
Some of the materials used include silicon and its oxides and nitrides, Pyrex, and various
metals and polymers. Apart from reducing the manufacturing costs due to mass
production, MEMS technology exploits some scaling benefits that allow physics-dictated
improvements in the operation of many systems upon downscaling. This technology can
have a significant impact on many fields such as the automotive industry, the electrical
appliance market, biotechnology, and healthcare.
The potential impact has sparked a
strong interest in, and funding of, MEMS research by many agencies to further develop this
technology.
A significant portion of the current MEMS research is centered on developing
integrated systems having many components that work in tandem. For example, some
Chapter 1. Introduction
components of a MEMS power converter may include a combustor, a generator, and
various pumps and valves. In many cases, an individual component with a specific function
is often required by different systems that perform different tasks. Because of this overlap,
small-scale devices are sought that can deliver certain functionalities currently lacking in
the MEMS field. One of those much-needed functionalities is the pumping of gases at high
flow rates.
1.2 Literature Review of MEMS Pumps
The majority of MEMS pumps that have been designed and/or demonstrated so far
handle liquids, not gases [2]. Liquid pumping finds many applications in the MEMS field,
such as fuel and propellant supply in microcombustors, microthrusters, and microengines
[3], diagnostic sampling of bodily fluids [1], drug delivery [1], and water management in
fuel cells [4]. Many MEMS liquid pumps have been reported in the literature. Some
examples include turbopumps [1], elastomeric fluid pumps [5], electro-osmotic fluid
pumps [6, 7], and a variety of reciprocating displacement pumps [8-13].
Gas pumping, on the other hand, has not received as much attention in MEMS
research. Most MEMS gas pumps that have been demonstrated so far fall into one of two
categories: displacement pumps or Knudsen pumps. Both of these kinds of pumps are
discussed below. There are also pumps that utilize the entrapment and adsorption of gas
particles, such as ion pumps. These pumps, however, consume the gas and are only good
for vacuum generation, not for other pumping applications.
addressed here.
Such pumps will not be
Chapter 1. Introduction
Displacement pumps operate by exerting pressure on the working fluid in a periodic
manner by means of a moving boundary [2]. In microscale pumps, this boundary is usually
a deformable plate or diaphragm with fixed edges, as opposed to a piston in macroscale
displacement pumps. The basic components of a displacement pump are the diaphragmbounded pump chamber, inlet and outlet check or actively actuated valves, and an
actuation mechanism. During operation, the externally-powered actuator displaces the
diaphragm, causing the chamber to periodically expand and contract. Fluid is sucked into
the chamber during expansion and then pushed out during contraction. The valves are
oriented (in the case of check valves) or actuated in a way to favor suction at the chamber
inlet and discharge at the outlet. Different diaphragm and valve actuation mechanisms have
been reported and/or proposed for MEMS displacement gas pumps, including
thermopneumatic [14], piezoelectric [15], electromagnetic [16], and electrostatic [17]
actuation.
Knudsen pumps operate based on the principle of thermal transpiration of rarefied
gases [18]. When two gas chambers are connected by a tube with a characteristic
dimension much smaller than (one tenth or less of) the mean free path of the gas, free
molecular flow exists. In this flow regime, the gas pressure and temperature ratios between
the two chambers are directly related, and this conclusion can be derived by balancing the
equilibrium molecular fluxes. As a result, varying the temperature between the two
chambers leads to a pressure differential which can be exploited for pumping. Knudsen
pumps are usually cascades of multiple, individually heated stages. Each stage has a
capillary section in which the gas is rarefied, and a connector section with larger
Chapter 1. Introduction
dimensions in which the flow approaches the continuum regime. A temperature increase is
imposed on the capillary section to produce a rise in the gas pressure by virtue of thermal
transpiration. The temperature is then reduced to its initial value in the connector section,
in which gas rarefaction is prevented due to the larger dimensions of this section. Due to
the absence of thermal transpiration in this section, the pressure drop is smaller in
magnitude than the pressure rise in the capillary section. This causes the gas to flow across
a net pressure increase in each stage, producing pumping action. Various micropumps
based on this operating principle have been reported in the literature [19-22].
Both kinds of gas pumps above suffer from certain disadvantages. The moving parts
in a displacement pump diminish its robustness and shorten its lifespan. Some moving
parts also require lubricants, which limit the use of these pumps in vacuum applications.
Since these pumps require many parts (actuator, valves, and chamber), their fabrication can
be challenging and dead volume can be an issue. Knudsen pumps do not contain any
moving parts, but they have other problems, such as their inefficient use of energy [18]. In
addition, Knudsen pumps' upper and lower pressures are usually limited, because the
mean free path of the gas has to be kept well above the dimensions of the capillary sections
and well below the dimensions of the connector sections for effective pumping. As the
pressure is reduced, the mean free path increases, and vice versa. In many cases, extending
each pressure limit separately requires functionalization and treatment of the pump's
internal surfaces. Increasing both limits simultaneously is usually challenging if possible at
all [18, 21].
For example, McNamara et al. [22] created a MEMS Knudsen pump for
evacuating a small cavity in a micromachined structure. The pump discharged into the
Chapter 1. Introduction
atmosphere, but the lowest pressure it was capable of creating in the cavity was 0.46 atm,
which is very high even for rough vacuum applications. Moreover, displacement and
Knudsen pumps have two major shortcomings in common: both require external power
sources for operation (to drive the actuator of a displacement pump and to heat the
capillary section of a Knudsen pump), which increase the overall pump size, and both have
low pumping capacities.
1.3 Ejector Pumps
1.3.1 Overview
Microscale ejector pumps offer a promising solution to the high flow rate gas
pumping requirement of many MEMS systems. Moreover, these pumps do not have the
drawbacks and limitations of the displacement and Knudsen pumps that were discussed
earlier.
1.3.2 Operating principle
The operation of an ejector pump is illustrated schematically in Figure 1.1. The
purpose of the system is to pump a fluid, referred to as the suction fluid, from a low
pressure Pio to a high pressure
Phigh.
To achieve this action, the pump mixes the suction
fluid with a high-speed stream referred to as the motive fluid. The motive fluid is
accelerated to a high speed before mixing to produce a low static pressure in the mixing
region. The low pressure causes the entrainment of the suction fluid. The mixing results in
momentum exchange: the suction fluid is accelerated and the motive fluid is slowed down
until both are moving at a common velocity. The combined flow is then slowed down by
....
....
....
....
: :: .: -::-...I_ :::
- Rer '-
Z- .._ .. _-qRR
..
- 1.
.. ..:.
. ... .-
-
_
..::.,:..:::::::.:
....
.......
:.,7 u : ':: .."
Chapter 1. Introduction
passing it through a de Laval nozzle. The nozzle allows pressure recovery by converting
part of the kinetic energy of the combined flow into pressure rise. As a result, the flow exits
at a pressure
Phigh
that is higher than Pion.
Suction
fluid at
pressure
Plow
de Laval nozzle
Combined
flow at
pressure
Phigh>
low
Figure 1.1. Schematic of an ejector pump
1.3.3 Macroscale characteristics and applications
Macroscale ejector pumps have been used historically for many purposes. Sample
applications include the mixing and compression of gases to obtain fuels with specific
heating values in petrochemical refineries, steam recirculation/recompression in the
evaporators of power plants and chemical and food processing systems, condensate
removal in heat exchangers, vacuum refrigeration and processing, and effluent pumping in
chemical lasers [23-25]. These pumps are attractive because they contain no moving parts
and require very little maintenance. They can operate at high efficiencies, and have been
used at the macroscale to produce a wide range of pressures, from high vacuum to aboveatmospheric [23].
Chapter 1. Introduction
1.3.4 Microscale ejector pumps
Microscale ejectors offer the same benefits of robust simplicity and no moving parts
as large-scale ejectors, plus one scale-dependent benefit. For a suction fluid at given
stagnation conditions, the mass flow rate that can be pumped through an ejector is, to first
order, proportional to the nozzle's throat area. The weight of the device, on the other hand,
scales with its volume. If a macroscale ejector is miniaturized while retaining the
proportions between its dimensions, the ratio of the pumping capacity to the system
weight scales inversely with length:
pumping capacity nozzle area
1
system weight
volume
length
Downscaling an ejector pump therefore allows for higher pumping capacity per unit
system weight. This is true as long as the system is not scaled down to the point at which
the thickness of boundary layers becomes comparable to the internal dimensions of the
pump. At that point, the boundary layers can cause significant reductions in mass flow rate
compared to the design value.
In microscale ejector pumps, the simplicity of design (e.g., no need for valves) and
the robustness due to the absence of moving parts and fatigue-inducing mechanical cycling
are significant advantages over displacement pumps. The potential for efficient operation
and wider pressure ranges favor ejectors over Knudsen pumps. Most importantly,
microscale ejector pumps are capable of handling higher mass flow rates per unit pump
volume than both displacement and Knudsen pumps, while achieving moderately high
pressure ratios. Multiple stages can be used to reach higher vacuum levels if desired. Table
1.1 compares the performance of some MEMS gas pumps in the literature, in terms of
20
Chapter 1. Introduction
pressure ratio and pumping capacity per device volume, to that of an ejector pump based
on the current work. The ejector pump specifications that are listed in the last row of this
table are derived in the analysis presented in Section 2.9.
Table 1.1. Comparison of the specifications of some MEMS gas pumps in the literature with
those of an ejector pump based on the current work
Maximum gas flow rate
(g/s/cm 3 of pump
volume)
Maximum
pressure
ratio
Research group
Pump type
Schomburg et aL [14]
Thermopneumatic
displacement pump
10
1:1.05
Kamper et aL [15]
Piezoelectric
displacement pump
10-4
1:2.8
Bohm et aL [16]
Electromagnetic
displacement pump
10
McNamara et aL [22]
Knudsen pump
1:2
Our group
Ejector pump
(predicted) based on
current work
1:10 per
stage
-7
-7
1:1.14
MEMS ejector pumps can thus potentially open up a wide range of opportunities for
MEMS applications if properly designed and demonstrated. Examples of these applications
include fuel and air feeds in microcombustors for power generation, and vacuum
maintenance in certain MEMS systems such as gas chromatographs. Another important
application is in pumping out the effluent of microscale chemical oxygen iodine laser
systems, as discussed in Section 1.4.1.
Chapter 1. Introduction
Very few microscale ejector pumps have been demonstrated so far. One challenge in
designing these pumps is providing a compact source for the motive fluid. Doms et al. [26]
and Chuech et al. [27] reported bulk-micromachined ejectors that were fabricated by
processing and bonding silicon and Pyrex wafers. The pump by Doms was tested with two
motive fluids: compressed nitrogen gas, and water vapor that was obtained by heating
liquid water outside the device. Chuech's pump was driven using air as the motive fluid.
Fan et al [28] reported an electro-discharge machined miniaturized stainless steel pump
with microscale features that used compressed air as the motive fluid. All of these pumps
were capable of achieving suction flow rates in the range 10-?-10-s kg/s, which are
relatively high compared to other MEMS gas pumps [2]. There are disadvantages, however,
associated with the motive fluids that were used to drive the reported ejector pumps.
Because of their low densities, gases such as air and nitrogen require large storage
volumes. This poses challenges for keeping the pump compact, especially in portable
systems. With pumps utilizing water vapor, more compact storage reservoirs are possible if
liquid water is stored inside the system and later vaporized during operation. This
approach, however, suffers from two significant drawbacks. One is the need for an external
power source to vaporize the water, which increases the size of the system. The second is
the integration of a vaporizing mechanism (such as an electric heater) into the system,
which can complicate the design and the fabrication process.
One approach that eliminates the needs for both large gas-storage reservoirs and
external power sources for liquid vaporization is the chemical generation of high-speed
gases from liquid reactants, usually via exothermic decomposition. For example, liquid
Chapter 1. Introduction
hydrazine (N2H 4) decomposes catalytically to produce ammonia, nitrogen, and hydrogen
gases at elevated temperatures. This decomposition has been widely employed in
propulsion applications to control and adjust the attitude of spacecraft, usually using an
iridium-alumina catalyst [29]. The gaseous products of the decomposition can be used, in
principle, to drive an ejector pump. Hydrazine's extreme toxicity and instability, however,
would pose serious safety and health challenges if this approach were implemented in
practice to generate the pump's motive fluid [30]. Storage is also a challenge because
hydrazine reacts with atmospheric carbon dioxide to produce corrosive compounds [31].
Other concerns are the toxicity and polluting capabilities of the generated ammonia [32,
33]. These problems translate into significant increases in costs for such systems [34].
1.4 Steam Generation from the Decomposition of High-Test
Hydrogen Peroxide
1.4.1 Overview and advantages
Another chemical reaction that allows the generation of a high-speed gas mixture
from a liquid precursor is the catalytic decomposition of high-test hydrogen peroxide, or
HTP. HTP refers to water-diluted hydrogen peroxide mixtures in which the peroxide mass
fraction is generally greater than about 70%. Hydrogen peroxide decomposes catalytically
to produce water, oxygen gas, and heat, according to the reaction:
1
2
H202 -+H20 +-02 + heat
(1.2)
Chapter 1. Introduction
When the initial peroxide concentration in a peroxide-water mixture is greater than about
67%, the energy released by the above decomposition in an adiabatic container originally
at room temperature is enough to vaporize all the liquid water present or produced. As a
result, a high-temperature mixture consisting predominantly of steam (and containing
some oxygen gas) can be generated. The mixture can then be passed through a nozzle to
convert some of its thermal energy to kinetic energy. This method can be used to produce a
high-speed flow suitable for driving an ejector pump.
The above approach offers the same benefits as using hydrazine, namely allowing a
compact container and not requiring any external power sources for vaporization, with
some advantageous differences. Unlike hydrazine, hydrogen peroxide is a nontoxic, or
"green", substance [31]. Despite posing some health hazards which are discussed in Section
4.2, hydrogen peroxide's low vapor pressure (1.95 mm Hg at room temperature [35])
prevents HTP mixture fumes from easily entering the human body. Exposure to peroxide
vapors emanating from an open HTP container at room temperature in a reasonably
ventilated area is not lethal to humans [31]. Also, hydrogen peroxide does not react with
the atmosphere, and can be stored for long periods of time in properly-designed
containers. The decomposition of peroxide according to (1.2) produces nontoxic,
environmentally-friendly products. All these features allow for increased safety and cost
reduction in peroxide systems compared to others that use more hazardous substances
[36].
Concentrated hydrogen peroxide has been used in many large-scale systems to
generate steam for driving ejector pumps. One such system is the Thiokol Hyprox, which
Chapter 1. Introduction
was successfully installed and used in many facilities to generate vacuum for rocket testing
[31]. A very promising potential application of an ejector pump that uses this approach at
the microscale is the aspiration of a micro-chemical oxygen iodine laser (pCOIL) system as
proposed by Wilhite et al. [37]. In this flowing-gas system, a 1.315 pIm wavelength laser
output is produced from the stimulated emission of excited atomic iodine, which is
obtained from the reaction of molecular iodine with singlet oxygen. The singlet oxygen is
generated from the multiphase reaction of basic hydrogen peroxide (a mixture of H202 and
KOH) and chlorine gas [38]. After laser emission, the low-pressure flow gas must be
discharged to the atmosphere to maintain stable operation. Since hydrogen peroxide is
already utilized in this application to generate the singlet oxygen, employing the above
ejector approach to pump out the flow gas is very desirable since it allows for a limitedinput portable p.COIL system.
1.4.2 Other examples and applications of high-test hydrogen peroxide
decomposition
Historically, highly-concentrated hydrogen peroxide has found its most widespread
use in propulsion applications [31]. HTP can be used as either a monopropellant or
bipropellant for rocket, torpedo, and submarine engines. As a monopropellant, hydrogen
peroxide is decomposed catalytically to produce a high-energy mixture of steam and
oxygen gas. This approach is very similar to that outlined in Section 1.4.1 for generating a
motive fluid for an ejector pump. The main difference is that in propulsion, the high-speed
exhaust mixture is used for thrust generation, whereas in an ejector the mixture is utilized
for momentum exchange. As a bipropellant, the peroxide is used as an oxidizer, in
25
Chapter 1. Introduction
conjunction with a fuel such as kerosene. The peroxide decomposes, either under the
action of a catalyst, or upon contact with the fuel itself, i.e. hypergolically, to produce
oxygen which causes the fuel to combust. The exhaust gases from the combustion are then
accelerated to provide thrust.
The first propulsion application of hydrogen peroxide dates back to 1933 Germany,
when Hellmuth Walter proposed using HTP mixtures at 80%-82% concentrations as
submarine monopropellants [31]. In the following years and during the World War II era,
the use of hydrogen peroxide expanded to include the propulsion of torpedoes and rocketassist devices for military aircraft. After the war, the technology moved from Germany to
the UK, USA, and USSR, and hydrogen peroxide received a great deal of interest. Some
famous applications included reaction control systems' thrusters (such as the X-1 and X-15
systems) that used peroxide as a monopropellant, and rocket engines (such as Black
Knight-Black Arrow, the AR engine series, and LR-40) that used peroxide as a bipropellant
oxidizer. This interest peaked during the 1950's and 1960's. Following this period,
peroxide was gradually displaced by chemicals with enhanced performance, such as
hydrazine in monopropellant applications and liquid oxygen and nitrogen tetroxide in
bipropellant applications. Ever since the 1990's, however, peroxide has seen a renewed
interest because of its previously-discussed advantages, such as its minimal environmental
impact, non-toxicity, and relative ease of handling. Many labs, agencies, and companies are
currently conducting heavy research on hydrogen peroxide-related systems. Examples of
these organizations include Lawrence Livermore National Laboratories, NASA, Rocketdyne,
Beal Aerospace, and Orbital Sciences Corporation [31].
Chapter 1. Introduction
Another important historical use of HTP has been as a gas generator for
turbopumps [31]. Peroxide is attractive for this application since the amount of water in
HTP mixtures can be controlled to produce decomposition temperatures below the
operational limits of the uncooled blades. One of the most famous systems that used a
peroxide-driven turbopump was the V-2 rocket which was developed toward the end of
World War II in Germany.
One interesting naturally-occurring system which uses peroxide as an oxidizer and
gas generator is the bombardier beetle [39]. As a defense mechanism, this beetle emits a jet
of noxious, boiling chemical spray that can be fatal to insects and small creatures. The
active components of this spray are benzoquinones produced from the oxidation of
hydroquinone by hydrogen peroxide. The beetle has two separate glands that enable this
mechanism: a large "reservoir" containing an aqueous hydrogen peroxide mixture (25%
peroxide by mass) and hydroquinone, and a smaller "reaction chamber" containing
enzymes such as catalase and peroxidase. When threatened, the beetle opens a musclecontrolled valve that forces the reservoir contents into the reaction chamber. The enzymes
catalyze the peroxide decomposition and the subsequent oxidation of the hydroquinone by
the generated oxygen gas. Both the decomposition and oxidation reactions are exothermic.
These reactions release enough heat to bring the mixture in the chamber to near its boiling
point, vaporizing about one fifth of it. Under the pressure of the vaporized gas and the
remaining (unused) oxygen from the peroxide decomposition, the mixture is expelled
explosively to the atmosphere through tips in the beetle's abdomen. The pressure causes
the entrance valve to close during this process, thereby protecting the insect's internal
Chapter 1. Introduction
organs. The defensive spray is ejected cyclically at about 500 pulses per second, and is
accompanied by audible "pop" sounds.
1.5 Thesis Roadmap
This thesis presents a MEMS device that generates steam (and oxygen gas) from the
catalytic decomposition of hydrogen peroxide. The device's effluent can be used to drive a
MEMS ejector pump. The different stages involved in developing the device are discussed,
including the modeling, design, fabrication, and testing.
In this chapter, the need for high flow rate MEMS gas pumping was discussed, and
ejector pumps were presented as an attractive option capable of delivering this
functionality. After describing the operation of these devices and comparing their
performance with other state-of-the-art pumps, the advantages of using hydrogen peroxide
to generate steam for driving ejectors were discussed. The chapter concluded with a
discussion of other examples in which the decomposition of high-test peroxide is utilized,
in industrial and military applications and even in nature.
Chapter 2 describes the design of the steam generator. After a discussion of the
challenges involved in designing this device, the conceptual design is presented along with
ways to counter those challenges. Detailed numerically-implemented physical models are
then developed to simulate the conditions in each of the different sections of the device.
The modeling results are used to design those sections. In addition, a thermal model is
developed to manage heat losses from the device, and it guides the design both of the
device and of a thermally-insulating package. The chapter also presents a parametric study
that was developed to examine the device operation under a range of conditions. The
28
Chapter 1. Introduction
chapter concludes with an analysis of an ejector pump based on the current device. This
analysis quantifies the proposed pump's improved performance over the state of the art.
Chapter 3 describes the fabrication of the device. The wafer-level die layout is first
discussed, along with the functions of the different die variations. The process flow of each
device layer is then presented, including cross sections of the layers after the main
processing steps, and the masks that were use for photolithography. The wafer bonding
process is then described. The chapter concludes with photographs of two diesawed
devices and cross-sectional SEM images.
Chapter 4 describes the task of setting up the test rig for experiments. The safety
concerns that accompany the use and storage of hydrogen peroxide are first discussed. The
machined package is then presented, and its material properties and design variations are
discussed. This discussion is accompanied by photographs of the original package and its
variations. The test rig components and overall setup are then described and displayed.
The chapter concludes with a discussion of the technique used to passivate those
components in order to make them ready for peroxide use.
Chapter 5 describes the experimental work. It starts with a brief discussion of the
experiments that were carried out using the mixer devices. The experiments on the
nominal devices are then presented. The operating conditions and the different
characterization methods, including visual inspection, refractometry, and temperature
measurements are described. The experimental results are then presented and used to
demonstrate successful operation. These findings are also compared with the modeling
results of Chapter 2 to provide comprehensive model verification.
Chapter 1. Introduction
Chapter 6 summarizes the presented research. Some of the challenges that were
encountered during the experiments are then described, and ways of alleviating those
challenges for improving future tests are suggested. The chapter concludes with a
discussion of some design modifications that will enable future use of the device in MEMS
pumping and micropropulsion applications.
CHAPTER TWO
2 MODELING AND DESIGN
2.1 Background and Challenges
Despite the many advantages of using hydrogen peroxide for steam generation, a
number of challenges exist, including the choice of catalyst, thermal management,
boundary layer effects, and safety considerations. For devices that generate steam via this
approach, thermal management and boundary layers are more important at the microscale
than at the macroscale.
The type of catalyst used can greatly affect the success of the peroxide
decomposition process. Typically, heterogeneous catalysts have been used in macroscale
decomposition devices [40-42]. These catalysts are usually solid materials such as silver or
manganese oxide. They are placed inside macroscale devices as meshed layers or pellet
beds, and inside microscale devices as channel wall coatings. Such static catalyst layers,
however, have a limited lifespan due to surface ablation upon repeated use [43]. The
lifetime of silver catalyst, for example, ranges from a few minutes to a maximum of about
30 minutes when using ultra pure peroxide [44]. Solid catalysts also necessitate the use of
rocket grade hydrogen peroxide which is relatively unstabilized. The use of stabilizers
allows longer-term storage and easier handling of the peroxide, but causes the poisoning of
solid catalysts. Furthermore, in microscale applications where very narrow, arrayed
channels are required to increase the surface area of the catalyst, flow clogging can become
a problem. Hitt et al. [45], in conjunction with the NASA Goddard Space Flight Center,
created a prototype monopropellant MEMS thruster to decompose hydrogen peroxide
Chapter 2. Modeling and design
using a heterogeneous catalyst. The device failed to achieve complete peroxide
decomposition and effluent vaporization. One of the causes to which this failure was
attributed was the formation of gas bubbles that may have blocked some of the reactor
channels. Using a homogeneous (liquid) catalyst [46] eliminates the limited lifespan issue
and allows the use of highly stabilized peroxide without causing poisoning. It also allows
wider reactor channels that are not as susceptible to clogging. To be able to use this kind of
catalyst, however, one extra challenge has to be met. For fast and complete peroxide
decomposition, the peroxide and catalyst streams must be well-mixed. This necessitates
the design and addition of a proper mixer section to the device. Mixing is especially tricky
at the microscale, where the laminar nature of flows usually makes diffusion the dominant
mixing mechanism [47], which significantly slows down the mixing.
Two further challenges arise from scaling considerations. The first challenge is
thermal management. The heat released from the peroxide decomposition is necessary to
keep the reaction going and to vaporize the water produced. The rate of heat generation
scales with the device volume (as in any homogeneous reaction). On the other hand, the
rate of heat loss from the device scales with surface area. If a macroscale device that
generates steam via the above approach is shrunk down while retaining the proportions
between its dimensions, the ratio of heat losses to heat generation thus scales inversely
with length:
heat loss
surface area
1
ocx
heat generation
volume
length
(2.1)
As seen in (2.1), downscaling makes the heat losses significant compared to the heat
generated by the reaction. When the scale gets small enough, the energy remaining inside
32
Chapter 2. Modeling and design
the device will become insufficient for sustaining the reaction and vaporizing the water
produced. This necessitates paying extra attention to thermal management to ensure
successful device operation. Osaki et al. [48] designed a microthruster that decomposes
hydrogen peroxide heterogeneously, and they found that the heat losses were so significant
that an electric heater was needed for effective operation. It is suspected that the increased
role of heat losses at the microscale may have also contributed to the failure of the device
by Hitt et al. That device lacked any thermal insulation.
The second scale-dependent challenge is boundary layer formation. The thickness 6
of boundary layers inside the device channels scales with the square root of distance x in
the flow direction. If a macroscale internal flow device is downscaled while retaining the
proportions between its dimensions, the ratio of the boundary layer thickness to the width
w of the flow channel thus increases:
w
oc
w
length
(2.2)
Downscaling can therefore cause boundary layers to occupy a large fraction of the flow
field, leading to a significant reduction in the effective flow area [49, 50]. In devices
designed to achieve supersonic flow, this reduction will prevent the flow from reaching the
sonic point if the scale gets small enough. This can be seen by examining the dependence of
Mach number M on the effective flow area A [51]:
dM2
2-
M2
1
dA
-2
1-NM A- .
(2.3)
As the flow approaches sonic conditions (M = 1), the denominator of the right-hand side of
(2.3) approaches zero. In this case, even very small area changes dA can cause large
33
Chapter 2. Modeling and design
deviations dM from the design Mach number. Therefore, proper compensation for the
boundary layer presence has to be incorporated into the device design.
One final challenge arises from practical considerations. Despite being nontoxic,
hydrogen peroxide is a strong oxidizer. This calls for careful safety provisions in any
experimental setup where high-test peroxide is used, to prevent explosions and
detonations.
In the current work, the catalyst, thermal management, and boundary layer
challenges were addressed by implementing multi-domain physical modeling. This
modeling was used to simulate the flow in different sections of the device and to evaluate
the heat losses from the device. The results guided the design of both a MEMS device that
decomposes hydrogen peroxide using a homogeneous catalyst and a package with
sufficiently high thermal resistance to enable full peroxide decomposition and complete
water vaporization. The model included the effects of boundary layers in the flow; these
effects were compensated for in the design. Finally, by using a continually-supplied
homogeneous liquid catalyst, the poisoning problem of heterogeneous catalysts was
eliminated, at the expense of adding a mixer section for the peroxide and catalyst streams.
Many safety measures for storing, handling, and experimenting with peroxide were taken
during the testing stage, as described in detail in Section 4.2.
2.2 Conceptual Design
In the current work, steam is generated from the decomposition of hydrogen
peroxide, and the reaction is facilitated by a homogeneous catalyst solution that is mixed
with the peroxide inside the device. The exothermic decomposition first produces oxygen
34
......................
..........m
.. ......
..
..
.................
. .......
__
__
. .
................................
..............................
. ..................
...
.......
....
Chapter 2. Modeling and design
gas and liquid water. The water is then vaporized by the heat generated, which leads to allgaseous products at the reactor exit. The products are subsequently accelerated to the
required speed, or expanded to the desired pressure. To enable these functionalities, the
device consists of three sections: a mixer for mixing the peroxide and catalyst streams, a
reactor for decomposing the peroxide and vaporizing the water, and a nozzle for
accelerating the gaseous products. A schematic of the device is shown in Figure 2.1.
Hydrogen peroxide reservoir
Reactor
Nozzle
Mixer
Catalyst reservoir
Figure 2.1. Conceptual design
The choice of catalyst solution was based on optimizing different performance
criteria, such as providing a strong and fast catalytic activity without containing large
insoluble particles that could clog the device channels [43, 46]. The ratio of catalyst to
peroxide flow rates was first guided by the findings of other groups that attempted to
optimize the reactor length required for full decomposition [52]. Preliminary experiments
were then conducted in which this ratio was slightly varied from the reported findings to
determine the value that lead to optimal operation. This optimal value was used in the final
device-characterization tests.
Chapter 2. Modeling and design
In the following sections, the models that were used to design the mixer, reactor,
nozzle, and thermally-insulating package are described. Even though these models are
presented sequentially, they are in fact interdependent, and iteration was required
between them to arrive at the final design.
2.3 Mixer Design
2.3.1 Mixer overview
The device includes a mixer section to enable the use of a homogeneous catalyst.
Homogeneous catalysts are supplied in solution form and make the catalysis process
volumetric. For uniformity and fast decomposition in the reactor, it is desired to achieve
thorough mixing of the peroxide and catalyst streams before the flow enters the reactor.
Passive micromixing is challenging since it tends to be dominated by diffusion due to the
laminar nature of microflows. This causes mixing to be slow and calls for long channels.
Active mixing can be faster, but it necessitates external energy sources to stir and mix the
flow.
In the current work, a fast passive mixer design (not more than a few millimeters in
length) was sought that would allow a compact device without requiring any external
energy sources for mixing. In addition, the mixer was designed to minimize the pressure
drop across it in order to avoid high supply pressures. This is necessary because the
interface between the device and the package consists of thin 0-ring seals, and a supply
pressure higher than 5 atm could lead to the failure of these seals [53]. All pressures listed
in this thesis are absolute pressures.
Chapter 2. Modeling and design
2.3.2 Micromixing literature review
Various groups have developed MEMS devices that can mix two or more flows. Some
of these designs achieve fast mixing by having out-of-plane (3D) features, such as the splitand-recombine design by Branebjerg et al. [54] and the serpentine design by Liu et al. [55].
In these designs, fast mixing comes at the expense of a complicated fabrication process and
a generally large pressure drop. Other designs have simpler fabrication, such as the Tshaped micromixer by Wong et al. [56], the Herringbone mixer by Stroock et al. [57], and
the impinging-jets mixer by Yang et al. [58]. These designs, however, suffer from other
problems when adapted to the current work. The Herringbone mixer and the standard Tshaped mixer are slow and require very long channels (with lengths on the order of
centimeters and meters, respectively) for sufficient mixing. The impinging jet mixer is fast
and compact enough (i.e. with a few millimeters long channel) only when one jet of
peroxide and one jet of catalyst are used. This makes the design not very robust.
Despite its slow mixing rate, the basic T-shaped mixer has an easily fabricatable
two-dimensional design that makes it very attractive from a MEMS-manufacturing point of
view. Methods for enhancing this mixer's performance were thus sought. Mengeaud et al.
[59] showed that zigzag channels provide better mixing than straight ones when the flow's
Reynolds number exceeds 80, due to laminar recirculations that are generated at the zigzag
angles. Wong et al. [60] showed that mixer walls with protrusions allow better mixing than
smooth walls, due to the generation of eddies and lateral velocity components as the flow
crosses these protrusions.
Chapter 2. Modeling and design
Furthermore, the Woias group [61, 62] showed that in a T-shaped micromixer, three
distinct laminar flow regimes can exist, each having different mixing characteristics. The
streamlines of the fluids to be mixed have different shapes at the mixer inlet for each of
these regimes. At very low flow speeds, there is the "stratified flow" regime in which the
streamlines are mostly straight and mixing is diffusion-dominated. At medium speeds,
there is the "vortex flow" regime in which vortices start building up inside the channels. In
this regime mixing is still primarily dominated by diffusion, but is slightly enhanced by the
swirling motion that drags fluid from the middle to the top and bottom of the mixing
channel. At high speeds, there is the "engulfment flow" regime in which the axial symmetry
of the flow breaks up and the streamlines interweave and reach to the respective opposite
half of the mixing channel. This causes a significant improvement in mass transfer and
allows this regime to have much faster mixing than the other two. Quantitatively, the three
regimes are distinguished by a dimensionless identification number K equal to the ratio of
the channel's hydraulic diameter dh,mixer to the Kolmogorov length scale Ak, which is the
scale of the smallest eddies in a turbulent flow:
K = dhmixer
(2.4)
/1k
Conceptually, K is a measure of the free space for the growth of vortices. At low values of K,
the fluid viscosity damps the starting of eddies, and mixing is slow, whereas at higher K
values, the flow conditions allow for the formation and growth of eddies and vortices that
enhance mixing. The Woias group has found in experiments and simulations that when K
exceeds a critical value of about 45, engulfment flow is achieved and mixing becomes
fastest.
Chapter 2. Modeling and design
2.3.3 Mixer modeling and design
The mixer was designed to have engulfment flow. By following the derivation in
[61], K can be expressed as:
K = lPdh,mixer Re3
pV2 Ly
(2.5)
Y.
In (2.5), Lv is the volume-to-area ratio of a control volume encompassing the mixer inlet
region as described in [61, 62], AP is the pressure drop in this control volume, p is the flow
density, V is the flow velocity, and Re is the Reynolds number based on
dh,mixer.
The ratio
AP/pV2 across the mixer inlet region can be estimated to first order by using the HagenPoiseuille equation:
AP
-
32
= ----
pV2
L(
(2.6)
Re dh,mixer
where Li is the equivalent length of the mixer inlet region, as described in [61, 62]. This
allows K to be expressed as:
%
K =32 L Re2
.
(2.7)
Ly
The Reynolds number can be determined using:
Re=
,4
ipdh,mixer
(2.8)
where rn is the mass flow rate and p is the flow viscosity.
For the mixer, it is desired to increase K to achieve engulfment flow while
minimizing the pressure drop to allow successful 0-ring sealing. Since each of these
requirements exhibits opposite dependencies on the mass flow rate as well as on the
39
Chapter 2. Modeling and design
channel's hydraulic diameter, an optimization study was carried out to find a combination
of flow rate and hydraulic diameter that meets the above criteria. The study was subject to
the constraint that the mass flow rate through the mixer has to be less than or equal to the
total flow rate across the device, which is fixed based on the pumping requirements
discussed in Section 2.9. A lower mixer mass flow rate simply means using multiple mixers
in parallel. Care was taken to ensure the device-to-mixer flow rate ratio produces a whole
number of mixers. The objective of the optimization was to find this number and the
dimensions of the mixers. Based on the results of this study, four identical mixers having
the design shown schematically in Figure 2.2 are used in parallel.
Peroxi
100 pm
Catalyst
Peroxide
30 pm x 40 pm protrusions
2.9 mm
Depth of channel and protrusions = 100 im
Figure 2.2. Schematic of a mixer with dimensions
Each mixer has three inlets: a middle catalyst branch sandwiched between two
peroxide branches. Compared to a two-inlet T mixer, this configuration cuts the distance
that the peroxide particles need to travel to meet the catalyst in half. In each mixer, the
Reynolds number is on the order of 500 under design conditions, and Li and Lv are found
following [61, 62] to be about 4 and 6 times the mixer width, respectively. This results in a
K value around 48, ensuring engulfment flow. Once the cross section is determined, the
length is chosen so that the residence time in the mixer (around 1 ms) matches that of
engulfment flow mixers reported in [62] to achieve high mixing qualities at their exit. This
40
Chapter 2. Modeling and design
time is 3 orders of magnitude less than the mixing timescale of a diffusive mixer with the
same width as the current design; the latter would have led to a much larger device. For the
original HTP concentration considered (see the discussion in Section 2.4.2 preceding Figure
2.4), the residence time in the mixer is less than the timescale for the onset of phase change
due to reaction. This was done on purpose to minimize the amount of bubbles forming
inside the mixer, which has narrower channels than any other section in the device, and
hence is the most susceptible to clogging. Ultimately, however, the device was operated at
higher peroxide mass fractions for which phase change begins in the mixer. This was not
found to cause any problems in operation, suggesting that some phase change in the mixer
can occur without causing adverse effects. To further enhance mixing following [59] and
[60], each mixer in the current design consists of 5 connected zigzag segments and has 12
wall protrusions that extend along the mixer depth.
The pressure loss across each mixer is estimated by first using (2.6) and replacing Li
by the total mixer length to determine the losses in a straight pipe of the same length and
cross section. Then the losses due to the bends are calculated by using the equivalent
length method [63], and those due to the protrusions are approximated by following the
experimental findings of [60]. This analysis shows that the total pressure drop in the
mixers is about 2 - 2.5 atm, which for a reactor pressure around 2 atm (see Section 2.4)
keeps the supply pressure below 5 atm as desired.
Chapter 2. Modeling and design
2.3.4 CFD simulation of the mixer
A numerical simulation of the mixer was carried out by Collins [64] using ADINA
software to study the mixing quality. The streams to be mixed were assigned different
colors, and non-reacting flow was assumed. The mixing quality was assessed based on the
color of the combined flow at the end of the mixer. The CFD analysis confirmed the design's
capability of achieving good mixing for the desired flow rates of peroxide and catalyst. The
simulation was two-dimensional due to computational limits, and therefore its pressure
calculations, which predicted much lower losses than the above analysis, were considered
inaccurate.
2.4 Reactor
2.4.1 Reactor overview
The reactor section is where most of the peroxide decomposition takes place. After
thorough mixing with the catalyst in the mixer, the peroxide starts decomposing in the
reactor according to (1.2). Along with the chemical reaction, other physical phenomena take
place inside the reactor. The hydrogen peroxide and water are initially in the liquid phase,
but are subsequently vaporized due to the heat released by the reaction. Some heat is also
lost through the reactor walls to the environment. The reactor has a rectangular crosssection for ease of microfabrication. The design process consists of determining the
dimensions of the reactor chamber that ensure complete peroxide decomposition and full
liquid vaporization. Although this may sound like a simple task at first, the intertwining of
physical phenomena occurring simultaneously inside the reactor makes the design
optimization challenging, since different phenomena point in different design directions.
42
Chapter 2. Modeling and design
For example, a larger reactor volume allows more residence time for the reaction to take
place, but also produces larger heat losses. A higher reactor pressure speeds up the
reaction after vaporization is complete, but requires higher peroxide and water boiling
temperatures, which can prevent the flow from ever reaching the vaporization stages if the
chamber is too short. The reactor model therefore had to account for all of the relevant
underlying physics, which necessitated a multi-domain modeling approach.
2.4.2 Reactor modeling and design
The reactor was designed to achieve complete peroxide decomposition and full
water vaporization. A qualitative understanding of what happens inside the reactor is
necessary for modeling, and is presented here first. Inside the reactor, hydrogen peroxide
decomposes according to (1.2) into H20 and oxygen while releasing heat. The flow in the
reactor passes through five stages based on the thermodynamic phases of the species
present, as summarized in Table 2.1. In stage 1, liquid peroxide decomposes into liquid
water and oxygen gas, and the heat released causes the reactor temperature to rise. Once
the boiling temperature of water at the reactor pressure is reached, stage 2 commences. In
this stage, liquid peroxide continues decomposing and the water starts vaporizing at
constant temperature. Once all of the water has been vaporized, stage 3 begins. In stage 3,
liquid peroxide decomposes to produce steam and oxygen gas, and the heat released again
causes the temperature to increase. This continues until the boiling point of the peroxide at
the reactor pressure is reached and stage 4 begins. In stage 4, the decomposition continues
while the peroxide changes phase at constant temperature. Once all of the peroxide has
been vaporized, stage 5 starts. In stage 5, the peroxide and H2 0 are both in the gas phase.
43
Chapter 2. Modeling and design
This stage continues until all of the peroxide has decomposed.
Table 2.1. Phases of H2 0 and H20 2 during the five reactor stages
H20
H20
2
Stage 1
Stage 2
Stage 3
liquid
liquid
liquid
liquid
liquid + gas
gas
Stage 4
liquid
+ gas
gas
Stage 5
gas
gas
To study the flow in the reactor, some approximations were made to allow
numerical modeling with reasonable time and computational effort. First, the models were
limited to steady-state behavior, and the experimental conditions were later chosen so that
the timescale of the transient effects was only a small fraction of each experimental run. A
bulk one-dimensional model was used, which is justified to first order by the reactor length
being about 4.4 times its hydraulic diameter in the final design. The accuracy of this onedimensional approximation was later verified experimentally. The reactor pressure was
assumed to be constant, since the two components of pressure loss (due to friction and
phase change) in the reactor are very small. The frictional pressure losses are negligible
compared to those occurring in the much narrower mixer channels. The phase-change
pressure losses were estimated following the homogeneous two-phase flow model in [65]
and were found to be negligible compared to the total pressure at the reactor inlet. Finally,
a very high heat transfer coefficient was assumed between the reacting particles and the
surrounding liquid during vaporization, which allows uniform-temperature modeling of
the flow during phase change. This last assumption is justified by the fact that the heat
44
Chapter 2. Modeling and design
transfer coefficient during liquid boiling is proportional to the latent heat of vaporization
[65], which is high for both peroxide and water. In addition, the phase change in this case is
volumetric, which leads to even higher heat transfer coefficients than with surface boiling.
All these observations indicate that the vaporization heat transfer coefficients are most
likely much larger than those describing other phenomena that are accounted for in the
current model, such as internal forced convection.
With the above assumptions, a state-space model was constructed in MATLAB. In
this model, the state or independent variables are taken to be the flow temperature T, the
peroxide mass fraction Yp, and the peroxide and water qualities,fp and fw, respectively. The
quality of a two-phase (liquid-gas) species is the mass of that species in the gas phase
divided by the total mass of that species present. The objective of the model was to study
the variation of these parameters versus distance x along the flow direction for different
reactor geometries, to determine which geometry results in full peroxide decomposition
(Yp = 0) and water vaporization (fw = 1) at the reactor exit. Table 2.2 describes the variation
of each state variable during the different stages. In this table, Tbw and Tbp are the boiling
points of water and peroxide respectively at the reactor pressure, and the last two rows
show the criteria used by the program to determine the beginning and end of each stage.
The stages are traversed sequentially.
Chapter 2. Modeling and design
Table 2.2. Variation of state variables during the different reactor stages
Stage 1
Stage 2
Stage 3
Stage 4
Stage 5
Yp
decreasing
decreasing
decreasing
decreasing
decreasing
T
increasing
TbW
increasing
Tb,p
increasing
fw
0
increasing
1
1
1
fp
0
0
0
increasing
1
Start
x=0
T=Tbw
fw= 1
T=Tbp
fp 1
End
T = Tw
fw=1
T = Tbp
fp=1
Yp=0
To study the evolution of the state variables versus distance x along the flow
direction (which is equivalent to time under steady-state conditions), the numerical
program divides the reactor into lengthwise differential elements. The mass, species, and
energy conservation principles are applied to these elements while accounting for the
chemical reaction within each element and the heat loss from the walls.
The conservation of mass simply states that the mass flow rate at any reactor
section has to be the same, since the reactor has one inlet and one exit. For the design
conditions, this total mass flow rate is 17.5x10-5 kg/s at any x.
The species law is used in conjunction with chemical kinetics to determine the rate
of change of peroxide mass fraction with distance. The rate constant k for the above
reaction is defined as the fraction of existing peroxide moles that will presently react per
unit time. Following [66] and [67], this rate constant is empirically given by:
Chapter 2. Modeling and design
k =10s
k=109exp
(Stage 1)
54800 kJ/kmol
(Stages 2-5)
(2.9)
where Ra is the universal gas constant. The rate constant expressions, however, are highly
dependent on the conditions of the experiments that were used to determine them. As a
result, the possibility that those expressions may not describe the decomposition in the
current case with full accuracy was considered. A conservative value for the package
thermal resistance was used to allow for some variation (about an order of magnitude) in k
without compromising the device's operation under design conditions. This is described in
more detail in Section 2.8. It should be noted, however, that all the simulation plots shown
in the thesis are based on the above expressions, for lack of better ones in the literature.
The reaction rate Rr, which is the number of moles of peroxide reacting per unit time and
per unit reactor volume, is then given by:
Rr
kp Yp
Mp
(2.10)
where My is the molecular weight of the peroxide. From the definition of Rr, the rate of
change of the peroxide mass fraction with distance in the flow direction is given by:
dYp
dx
MpRrA
(2.11)
i
where A is the effective flow area, which for the reactor is equal to the chamber's crosssectional area.
Chapter 2. Modeling and design
To satisfy the conservation of energy, the first law of thermodynamics for steadystate open systems with no shaft work is applied to the differential reactor element in
Figure 2.3:
rhh-Lrhh
hhax] -IOSS =0.
rhh ---
Reaction
4
-
Ax
--
(2.12)
rhh+
(rhhjax
-
Figure 2.3. Differential reactor element
In (2.12), h is the enthalpy of the flow and Q10s, is the rate of heat loss from the walls. The
flow enthalpy is the mass-weighted sum of the individual enthalpies hP of peroxide in any
phase, hw of H20 in any phase, and ho of oxygen gas:
h=hpYp +hwYw +hOYO.
(2.13)
In (2.13), Yw and Yo are the mass fractions of H2 0 and oxygen respectively, and are
determined from stoichiometry and the initial peroxide mass fraction Ypi:
Yw =(1 -Yi
)+ Mw (YP'i -Ye ,
MP
Yo = 0 5 M (YPi -Y )MP
(2.14)
(2.15)
In (2.14) and (2.15), Mw and Mo are the molecular weights of H2 0 and oxygen respectively.
The individual enthalpies in (2.13) are given by:
Chapter 2. Modeling and design
hp= hfp(T) + fphfg,p(T),
(2.16)
hw= hfw (T) + fwhfg,w(T ),
(2.17)
ho = ho (T).
(2.18)
In (2.16)-(2.18), hf and hfg denote liquid-phase enthalpy and heat of vaporization,
respectively. All temperature dependent variables in (2.16)-(2.18) are determined from
property databases [35] and account for the heat of formation of each species. With some
algebraic manipulation, (2.13)-(2.18) can be used to express (2.12) in the form:
- rhAX c -- +fYphfP
+Yw
dx
dx +
f
dfg,w
dYp =$sR
dx
Mp dx )10ss
(2.19)
In (2.19), c, is the mass-weighted average specific heat of the flow at constant pressure
(equal to the partial derivative of the flow enthalpy with respect to temperature), and AHR
is the heat of reaction, defined as:
AHR =-hpMp +hwMw +O.5hOMo.
(2.20)
The heat of reaction describes the amount of heat released when 1 mole of peroxide
decomposes. Note that AHR as defined here is a function of temperature and H20 and
peroxide qualities. A representative value typically quoted in the literature is AH" = -96
kJ/mol; this is the standard heat of reaction for the decomposition of liquid peroxide at
250C into liquid water and oxygen gas. The negative sign denotes an exothermic reaction.
The flow loses heat mainly by internal forced convection to the reactor walls. Heat
transfer by radiation was estimated and found to be negligible compared to convection.
The heat loss term in (2.12) is therefore given by:
Chapter 2. Modeling and design
$10ss
= 2h1c (Wreactor + hreactor)Ax(T
- Twaji),
(2.21)
where Wreactor and hreactor are the reactor width and depth respectively, Twaji is the device
wall temperature (determined by iteration as described in Section 2.8), and hic is the
internal convection coefficient, given by:
hic = Nuic d kcond
(2.22)
h,reactor
In (2.22), Nuic is the Nusselt number for internal convection, kcond is the mass-weighted
average thermal conductivity of the flow, and dhreactor is the hydraulic diameter of the
reactor. The thermal conductivities of the individual species are determined from property
tables [35], and the Nusselt number is found to be in the range 25-75 from empirical
internal flow correlations that account for entry effects [65]. Finally, (2.11) and (2.21) can
be used to express (2.19) as:
c(
+YThg,P
+Yw h
w
-AH
A - 2hc (wreactor + hreactor TTwa ).
(2.23)
Note that two of the three terms inside parentheses on the left-hand side of (2.23) vanish
for each stage in Table 2.2. Therefore (2.11) and (2.23) constitute a state-space system of
coupled differential equations. This system is solved in MATLAB using the odel5s solver
which is capable of handling stiff differential equations. To determine the onset of stages 2
and 4, the flow temperature at every x is compared to the saturation temperatures of water
and peroxide, respectively, at the reactor pressure. This pressure, for operation at the
design point, is set at around 2 atm in order to keep the device supply pressure less than 5
atm while allowing for the mixer pressure losses described earlier. Expressions for these
saturation temperatures as functions of pressure are also obtained from property tables
50
Chapter 2. Modeling and design
[35]. Once the state variables are evaluated, the flow density p can be estimated using
constitutive relations, such as the ideal gas law in stage 5:
P
p = reactor
RT
(2.24)
In (2.24), Preactor is the reactor pressure and R is the mass-based gas constant of the flow.
The velocity V can then be found using
V
pA
(2.25)
Using the above model, the reactor was designed to be a 15 mm long rectangular
chamber with a 4mm x 3mm cross section. The flow temperature and species mass
fractions that are predicted by the model are plotted versus distance along the reactor in
the flow direction in Figures 2.4 and 2.5 respectively. These figures are for the case of a
90% HTP mixture supplied at 7.0 mL/min and catalyzed by a ferrous chloride tetrahydrate
solution (80% saturated) at 0.5 mL/min, which corresponds to an overall initial peroxide
mass fraction of 83% (after mixing with the catalyst). These conditions will be referred to
as the design conditions since the device performance, in terms of the energy content of the
exiting flow, is best under those conditions. The device, however, was originally designed
for more conservative operation with an 80% HTP mixture. The 80% concentration,
although more desirable than 90% from a safety point of view, was not commercially
available. This explains the extra reactor length beyond the point of complete peroxide
decomposition for the 90% HTP case, as seen in Figure 2.5. The extra length is also visible
as the region in Figure 2.4 where the temperature starts to drop after peaking. The drop is
.......................................................................................
............
. .....
...
...............
W%V%
R-
Chapter 2. Modeling and design
due to heat losses from the device walls after the decomposition is complete and the heat
generation stops. The inserts in Figure 2.4 show the vaporization stages.
900
850800-5
750-
S700-
445-
650
435
60 E50
0
0.1
0.2
0.3
550-----------------.395
.*500-,
450
394
400
393
350
0.010--0
0.02 -0 .03,
300
0
1
2
3
4
5
6
7
8
9 10 11 12 13 14 15
Distance from reactor inlet along flow direction (mm)
Figure 2.4. Plot of the flow temperature along the reactor length for the design conditions,
with inserts magnifying the vaporization stages
0.9
-0.8
0.8
-
10.7
0I
0.7
H20
"" " - - " " - "- - - - - - - -0.-----
C 0.
r-0.0
-0.-............-----.--2...-.
0.4
0.
.A
0.2
0.1
2
0.6 -- -
0
\
--
1 0.3
0.2
I2
H20
2
0.1
0
0 1 2 3 4 5 6 7 8 9 101112131415
Distance from reactor inlet along flow direction (mm)
00
- 0
0.1
0
-.
0.2
2
0.3/
Figure 2.5. Plot of the species' mass fractions along the reactor length for the design
conditions, with an insert magnifying the region near the reactor inlet
Chapter 2. Modeling and design
2.5 Nozzle
2.5.1 Nozzle overview
The nozzle accelerates the flow from the relatively low reactor speeds to supersonic
speeds, thereby expanding it to a pressure lower than that inside the reactor. When the
steam generator is used to drive an ejector pump, the nozzle should be designed so that its
exit pressure matches the ejector's upstream pressure. In the current work, the device was
tested for proof-of-concept only, and the flow exited into the atmosphere. The nozzle was
therefore designed to expand the flow into atmospheric pressure. This prevents the
formation of shock waves at the exit, which would dissipate part of the flow energy. To
achieve this expansion, the nozzle requires two sections: a converging section in which the
flow is accelerated to sonic velocity, and a diverging section in which the flow becomes
supersonic. The nozzle design process consists of sizing the different nozzle sections to
achieve the desired expansion.
2.5.2 Nozzle modeling and design
An isentropic one-dimensional compressible flow model was first used to
approximate the dimensions of the throat and exit sections, and then the design was
refined by accounting for heat losses and boundary layer formation. The nozzle depth was
kept the same as that of the reactor (3 mm) to simplify the device fabrication.
Figure 2.6 is a nozzle schematic showing the known quantities at different locations.
The isentropic model uses these known quantities to calculate the throat and exit areas as
described below.
Chapter 2. Modeling and design
Nozzle inlet
Nozzle throat
Nozzle exit
Area and all
flow variables
are known
Flow is sonic
(Mach number = 1)
Pressure is
atmospheric
Figure 2.6. Schematic of the nozzle showing the known quantities at different locations
At any location along the nozzle, the Mach number can be determined from the flow
velocity and temperature, using
M=
(2.26)
In (2.26), y is the specific heat ratio of the flow in the nozzle, which is assumed to be
composed of oxygen and steam in the concentrations given by the reactor model. This ratio
is determined from the specific heat and gas constant of the flow, which are calculated as
the mass-weighted sums of the individual respective properties of steam and oxygen. The
isentropic M-A relation is then applied between the inlet section and the throat section to
calculate the necessary throat area Ath, using (2.26) to estimate the Mach number Mi at the
Chapter 2. Modeling and design
nozzle inlet. This inlet has a cross-sectional area At equal to that of the reactor, and the flow
is assumed to be sonic (Mth = 1) at the throat. This results in the expression:
A
'
1+
2
0.5(y+1)
1(2.27)
0.5(y+l)
-A th
1+
M
'
2
To determine the nozzle exit area, the necessary exit Mach number Me is first calculated by
applying the M-P isentropic relation between the nozzle inlet and exit sections, using:
Y
Y-1
+ Y-1 Me2
Pi
Pe
__2
(2.28)
_
1+ Y-1
2
M2
J
In (2.28), Pi is the known inlet pressure (i.e. reactor pressure) and Pe is the exit pressure,
which is deliberately matched to that of the atmosphere to prevent the formation of shock
or expansion waves at the nozzle exit. Then the isentropic M-A relation is applied again,
this time between the inlet and exit sections, to calculate the exit area Ae:
A
'i O.S(y+l)-=A e
1 + Y-Me
1+ Y-M2
21~YIM;
e
K-
21Y1~ e
0.5(y+l)
((2.29)
-
-
The nozzle length is set somewhat arbitrarily at 2.3 mm by trying to minimize the
overall device volume without causing strongly non-ideal flow behavior, such as separation
from the nozzle walls. The isentropic flow assumption is then relaxed, and the variations of
the flow Mach number M, stagnation temperature Tt, and stagnation pressure Pt with
distance x along the flow direction are determined by using a non-isentropic compressible
flow model [51]. This model accounts for heat loss from the nozzle walls but assumes that
*._
.
....
........
:
..........
-111,
..........
.......
..........
...............
Chapter 2. Modeling and design
the flow has constant molecular weight and specific heat. The nozzle is divided into
lengthwise differential elements as shown in Figure 2.7, and the conservation of mass,
energy, and momentum are applied to each element.
Heat losses
A+dA,
A,
FAw
Pt
M+dM,
Pt+dPt
Figure 2.7. Differential nozzle element
By following the derivation in [51], the following equations are obtained:
dPt
dx
_
d__z___M(
dM 2
m2
1+
2
1-M
dx
IC
__
dPt
dx
2
nozzle + hnozzle XT ~
)wall
'Y2dA 1+yMz dTe'
M 2 ----+
,d
(}Adx
Tt
dx)
- yM 2 Pt dTt
dx
2Tt
(2.30)
(2.31)
(2.32)
In (2.30), hic is determined using (2.22) after replacing dAreactor with dhnozzle, the nozzle's
hydraulic diameter. Empirical correlations [65] are used to calculate Nuic, which is found to
be in the range 480-560 for the nozzle. The higher nozzle speeds result in higher Nusselt
numbers compared to the reactor. The above state-space system (2.30)-(2.32) is then
solved in MATLAB using the ode1Ss solver. Once the stagnation properties are found, the
static flow temperature T and pressure P at any x can then be determined using:
Chapter 2. Modeling and design
=1 +
m2,
2
T
=
P
1
m2
Y
(2.33)
(2.34)
2
The flow density and velocity along the nozzle are calculated using (2.24) and (2.25)
respectively, and are used to estimate the thickness of the boundary layers in the nozzle.
These layers tend to lower the mass flow rate from its ideal (frictionless) value. Their effect
is most pronounced near the nozzle throat, where they could prevent the design from
achieving sonic (and subsequently supersonic) conditions unless accounted for. The
displacement thickness is selected as the measure of the boundary layer thickness, and is
defined as the distance by which the walls in a viscous flow would have to be pulled apart
to maintain the same mass flow rate as a hypothetical frictionless flow with the same
density and initial wall separation. The displacement thickness 6*is calculated following the
Blasius solution [68]:
6*2
=21.72
IriX.
(2.35)
In (2.35), xn is distance measured from the nozzle's inlet along the flow direction inside the
nozzle, and p is the flow viscosity which is obtained as a function of temperature from
property tables [35]. The factor of 2 in (2.35) accounts for the formation of boundary layers
on both lateral sidewalls of the nozzle; the effects of top and bottom wall boundary layers
are less critical since the nozzle width is much smaller than its depth near the throat. Thus
the nozzle width must be increased by 6* to leave the mass flow rate unaltered. Figure 2.8
plots the nozzle profile before and after boundary layer compensation, with an insert
showing the profile in the vicinity of the throat. Note that the actual fabricated design
57
Chapter 2. Modeling and design
replaces the sharp corners near the throat with filleted ones to lower the frictional losses.
The fillet radius is about 200 prm, and the throat section is made slightly longer (by about
the same value) to ensure that the throat width given by the model is met in the design.
2
After b.. compensation
1.5
- - - - Before b.i. compensation
E
-0.1
2 0.5
2
0
0
-0.5 N
o
Z
-1-
-0.1
:-0.2
1.7
-1.5
- - - - -
-0 .
1.75
1.8
1.85
-...............
-2
0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 2 2.2
Distance from nozzle inlet along flow direction (mm)
Figure 2.8. Nozzle width profile before and after boundary layer compensation with insert
showing the throat region
Figure 2.9 plots the variation of static and stagnation pressures along the nozzle for
the design flow conditions. The static pressure at the exit is just above atmospheric
pressure, which prevents the formation of shockwaves at that section, as desired. Figure
2.10 plots the Mach number variation along the nozzle for the design conditions. The
apparent discontinuity in the slopes of the curves in these two figures near the throat
section is due to the computational limits of the numerical program and has no physical
significance.
Chapter 2. Modeling and design
1.8 CD
-
1.6 -
Static pressure
Stagnation pressure
S1.4
CD
oL 1.2
0
0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 2 2.2
Distance from nozzle inlet along flow direction (mm)
Figure 2.9. Plot of static and stagnation pressures along the nozzle for the design
conditions
1.2
1
.0 0.8
E
0.6
0.4
0.2
0
0 0.2 0.4 0.6 0.8
1 1.2 1.4 1.6 1.8
2 2.2
Distance from nozzle inlet along flow direction (mm)
Figure 2.10. Plot of the Mach number along the nozzle for the design conditions
2.6 Parametric Study
In addition to simulating the design conditions, the numerical programs for
modeling the reactor and nozzle were expanded to study the device operation over a range
of overall initial peroxide mass fractions between 71% and 83%. This parametric study
was used to assess the accuracy of the models by comparing them to the experimental
results discussed in Section 5.3.
Chapter 2. Modeling and design
Figure 2.11 plots the mass fraction of peroxide exiting the reactor versus the initial
peroxide
mass fraction. Full peroxide decomposition occurs
for overall initial
concentrations down to about 74%, below which some peroxide starts exiting the device
undecomposed. Conceptually, this is due to the decreasing energy content of the HTP
mixture as the peroxide concentration decreases. This causes the temperature-dependent
reaction rate to slow down until the residence time of the flow in the device is no longer
enough to decompose all of the peroxide. Even though decomposition is incomplete for all
initial concentrations less than 74%, the part of the curve in Figure 2.11 between initial
concentrations of 71% and 73% is not monotonic due to two competing effects that
influence the flow temperature in opposite ways. As the initial peroxide concentration is
lowered, the energy content of the mixture decreases, but also less heat is lost from the
device walls to the environment due to the lower wall temperatures reached.
5.5
5
S 4.5 -0
CU
(D~
(
3.5-
S3-
-
2.5-
E~
2
V
1.5-
2
1
-0.50
71 72 73 74 75 76 77 78 79 80 81
O\erall initial peroxide mass fraction (%)
82
83
Figure 2.11. Plot of peroxide mass fraction at the reactor exit versus initial peroxide mass
fraction
Figure 2.12 plots the variation of the static and stagnation temperatures at the
nozzle exit versus initial peroxide mass fraction. The shapes of these curves are directly
60
Chapter 2. Modeling and design
linked to the degree of peroxide decomposition in the reactor from Figure 2.11. To explain
the shapes of the temperature curves, an approximate steady-state energy balance is
performed on the entire reactor as shown schematically in Figure 2.13. This balance can be
expressed as:
Rate of chemicalenergy generatedinside reactor=
Rate of thermalenergy out - Rate of thermal energy in + Rate of heat loss.
(2.36)
750
Stagnation temperature
Static temperature
700 ------
650
-
600
',.Po-
550
,I
500
/I
450
.
-
400
-
350
71
72
73
74
75
76
77
78
79
80
81
82
83
Overall initial peroxide mass fraction (%)
Figure 2.12. Plot of static and stagnation temperatures at nozzle exit versus initial
peroxide mass fraction
Heat lost to environment
Thermal energy in---+ Chemical energy generated inside reactor --
Thermal energy out
Figure 2.13. Energy balance on the entire reactor
The flow energies in and out of the control volume only include the thermal
components of the enthalpies, since the chemical components are accounted for as energy
Chapter 2. Modeling and design
generation inside the reactor. By assuming an average and constant specific heat e, for the
flow, the difference in these thermal flow energies can be expressed as rn,(Tr,ex - Tr,in),
where Tr, in and Tr, ex are the flow temperatures at the reactor inlet and exit, respectively.
The heat loss term in (2.36) is proportional to the temperature difference between the
reactor wall and the environment. Since higher flow temperatures at the exit are correlated
with higher wall temperatures, it is reasonable to assume a linear dependence between
these two temperatures, to first order and for the purpose of this approximate analysis
only. Then the right-hand side of (2.36) can be expressed as:
Rate of thermal energy out - Rate of thermal energy in + Rate of heat loss =
C(2.37)
w ear~,ex ++he
C*,
where a is a proportionality constant, and C*is the sum of any remaining constant terms,
including those that contain Tr, in.
When the peroxide decomposition is complete, which from Figure 2.11 is observed
for initial peroxide mass fractions greater than 74%, the energy generation term in (2.36)
can be expressed as:
Rate of chemical energy generatedinside reactor= (rh, / M, IAHIR ,
(2.38)
where rn, is the mass flow rate of peroxide in the supply mixture, and AHR is an average
constant heat of reaction. When (2.37) and (2.38) are substituted in (2.36), it becomes clear
that the flow temperature at the exit of the reactor is linearly proportional to the ratio of
peroxide to total mass flow rates in the supply mixture, which by definition is the peroxide
initial mass fraction, for the range over which the peroxide decomposition is complete. This
exit temperature closely matches the nozzle inlet stagnation temperature due to the low
62
Chapter 2. Modeling and design
flow speed in the reactor. In the nozzle, the exit static and stagnation temperatures follow
the dependence of this inlet stagnation temperature, producing the linear portions of the
curves in Figure 2.12.
When the peroxide decomposition is incomplete, the energy generation term is no
longer proportional to the initial peroxide mass fraction, since some peroxide now leaves
the reactor undecomposed. Therefore the linear dependence of the temperatures on initial
peroxide concentration breaks, producing the nonlinear portions of the curves in Figure
2.12.
Since the same nozzle design is used for the entire range of the parametric study, it
is expected that the pressure in the reactor chamber will vary with initial peroxide
concentration. This variation can be explained by examining the relationship between
mass flow rate, Mach number, and stagnation temperature and pressure in a compressible
flow:
AP~
J
YMj(2.39)
2
-1
The throat Mach number is ideally unity, so the right-hand side of (2.39) is always a
constant quantity at the throat. Moreover, the mass flow rate and throat area are the same
for the entire parametric study. Assuming constant flow properties, the stagnation
pressure at the throat will vary as the square root of the stagnation temperature, which is
close to a linear variation due to the small range involved. If the nozzle's viscous and heat
losses are neglected for the sake of this argument only, isentropic flow ensues, and the
stagnation pressure and temperature are uniform along the length of the nozzle. Therefore,
63
Chapter 2. Modeling and design
as the stagnation temperature of the flow decreases with the initial peroxide concentration,
the stagnation pressure must also decrease proportionally to keep the other quantities in
(2.39) fixed. The reactor pressure closely matches the nozzle stagnation pressure due to the
low flow speed inside the reactor. As a result, the reactor pressure follows a trend similar
to the temperature curves in Figure 2.12. Figure 2.14 plots the variation of the reactor
pressure over the above range of initial peroxide mass fractions.
E 1.9-
1.8-
I 1.7 0
CU)
5 1.61.5
71
'1
72 73
74 75
76
77
78
79 80
81
822
8
83
Overall initial peroxide mass fraction (%)
Figure 2.14. Plot of reactor pressure versus initial peroxide mass fraction
The reactor pressure for the design conditions is determined from the maximum
supply pressure requirement discussed earlier, and the nozzle is designed accordingly to
produce atmospheric static pressure at the exit. For the other operating conditions in the
parametric study, the reactor pressure is determined by iteration using both the reactor
and nozzle models. First a given pressure is assumed, and the reactor model is solved to
determine the conditions at the reactor exit, which serve as boundary conditions for the
nozzle. The nozzle model is then solved to determine the conditions along the nozzle. By
iteration, the reactor pressure is selected as the value that leads to sonic flow (M = 1) at the
throat.
.........
......
.
............
...............
Chapter 2. Modeling and design
2.7 Overall Design
The design of the overall device is shown in Figure 2.15. The top layer is made of
Pyrex as discussed in Chapter 3, and it allows optical access into the device's interior. The
reactor, nozzle, and one of the four parallel mixers can be seen. In addition, there are two
sets of through holes that terminate just before the Pyrex layer: two large holes (1 mm
diameter) for connecting the peroxide and catalyst supplies, and six smaller holes (0.5 mm
diameter) for inserting thermocouples to measure the device wall temperature during
experiments. The openings of all these holes are on the back side of the device and are not
visible in Figure 2.15. Finally, a buffer zone is left intentionally between the nozzle exit and
the device edge to protect the nozzle during diesawing.
Nozzle
Reactor
Mixer
Thermiocouple
insertion holes
Catalyst inlet hole
Peroxide inlet hole
Figure 2.15. 3D model of entire device
Chapter 2. Modeling and design
2.8 Thermal Management
Thermal management is key to ensuring successful operation of the device. The
right-hand side of (2.23) shows that there is competition between the heat generation
inside the device and the heat loss from the walls. The ratio of heat losses to heat
generation increases as the device is downscaled, as was shown in (2.1). Without proper
management, the heat loss term can dominate, preventing the flow temperature inside the
device from rising to the desired values.
This can cause incomplete peroxide
decomposition by slowing down the temperature-dependent reaction rate constant in
stage 5, and can also lead to incomplete water vaporization. To minimize the heat losses,
two packaging schemes were evaluated.
The first approach consisted of suspending the device in air and making the fluidic
connections using low thermal conductance pipes adhered directly to the device surface
(e.g. by epoxy). In this case, the dominant heat loss mechanism is internal forced convection
from the flow to the device walls followed by external natural convection from the walls to
the ambient air. The coefficient of natural convection hNC from each side of the device to the
ambient air is given by:
hNC = NuNC
k.
ir
(2.40)
where kair is the thermal conductivity of air, Lside is the characteristic length of each side,
and NuNC is the Nusselt number for external natural convection, found using empirical
correlations [65]. The convection coefficients were found to be in the range 15-30 W/m 2 K,
providing sufficient thermal insulation for complete peroxide decomposition and full water
vaporization. This approach, however, was ultimately not selected due to robustness
66
Chapter 2. Modeling and design
considerations: connections by direct adhesion to the device are prone to leakage and
failure, especially for the somewhat high supply pressures anticipated (4-5 atm).
The second approach consisted of encasing the device inside a machined package
using O-rings at the interface, with threaded fittings connecting the supply pipes to
threaded ports in the package. In this approach, the flow primarily loses heat by internal
forced convection to the device walls and then by conduction through the package and
natural convection from any exposed device areas not covered by the package to the
environment. This is shown schematically in Figure 2.16. This approach was selected
because it was more robust than the first. The package material was chosen to provide
sufficient thermal insulation as described below, in addition to meeting other
requirements, such as having a melting point higher than about 4000C to tolerate the
reactor wall temperatures, having a machining tolerance at least as tight as 0.05 mm to
allow successful installation of the 0-rings, and being compatible with peroxide.
To determine the minimum required thermal resistance of the package, the
following algorithm was used. First the resistance due to conduction within the silicon
walls was estimated and found to be about an order of magnitude less than that due to the
flow's internal convection. The wall resistance was thus neglected, and the walls were
assumed to be isothermal. Then the thermal network shown in Figure 2.17 was used to
study the heat transfer in Figure 2.16 at steady state. In Figure 2.17,
resistance of the package, Rcon,
m is
Rpack
is the thermal
the convective thermal resistance in each differential
element of the reactor (Figure 2.3) or nozzle (Figure 2.7) at an average element
temperature Tm, and Too is the ambient temperature.
.....
.
... ...
....
Chapter 2. Modeling and design
Environment at ambient
Natural convection
from exposed device
walls to environment
temperature Too
External package
surface at ambient
temperature Too
Device walls at
temperature Twai
Figure 2.16. Schematic of a cross-section of the device at distance x along the flow
direction, illustrating the dominant heat transfer mechanisms for the configuration in
which the device is encased inside a package
Rpack
Twall
T3
.. ...
. Tn
Figure 2.17. Thermal network between device and environment at steady state
Chapter 2. Modeling and design
The convective resistance of each element is given by:
1
Rcon,
(2.41)
AAsurf
where hic is given by (2.22) and AAsurf is the differential surface area over which convection
takes place. The program finds Twan by iteration, starting with a guess value to calculate the
node temperatures Tm using the reactor and nozzle models, and then using these node
temperatures to refine the value of Twain by applying Kirchoff's current law:
n
wa
Rcony,m
Twal
_R
To"
(2.42)
pack
This process is repeated until Twain converges. Note that in (2.42)
Rpack is
not known yet, so
different values are tried with the reactor and nozzle models to determine the minimum
resistance that will allow complete peroxide decomposition and water vaporization, with
no condensation in the nozzle. This minimum value is found to be in the range 0.5 - 4 K/W
and is used to determine the maximum allowable package thermal conductivity
kpack
for a
given geometry, such as the one shown in Figure 2.18. The above range in the required
thermal resistance is mainly due to the uncertainty in the reaction rate constant k. The
upper end of the range allows for a decrease of about an order of magnitude in k from the
values given in (2.9). The outer package surfaces are assumed to be at the ambient
temperature, and the device surfaces that are directly exposed to air are cooled by natural
convection with negligible contribution from radiation. The package resistance is thus
estimated to first order using a one-dimensional heat transfer analysis:
Chapter 2. Modeling and design
Rpack
=
kpacA +hNC ANC
(2.43)
In (2.43) A; and L; are the cross sectional area and length of each conduction heat path in
the package between the device and the environment, ANC is the device surface area that is
directly exposed to air, and
hNC
is estimated using (2.40) and empirical correlations for NuNC
[65]. This analysis shows that the package thermal conductivity has to be no larger than
0.8-3.2 W/m-K. To allow for a conservative design, a value of 0.4 W/m-K is used in the
simulations. The thermal conductivity of the material used for constructing the actual
package also has this value, as will be described in Section 4.3.
The package is shown schematically in Figure 2.18. It consists of two parts that are
clamped together using fasteners. The top part has a slot in which the device sits and a
window for optical access during the experiments. The bottom part has holes for
thermocouple insertion, 0-ring glands, and threaded ports on the back side (not visible in
Figure 2.18) for connecting the peroxide and catalyst supplies. The detailed design of the 0ring glands is shown in Figure 2.19.
..............
1.11,
.............................................
.
Chapter 2. Modeling and design
Optical
access
window
12. Omm
Threaded
holes for
fasteners
Device
slot
Figure 2.18. Package design
*,1
/:fI(
I IF/7T::
p1.50
+1
//
0
0
C0
I--%
1.40
0. 5
7 +0.05
Figure 2.19. 0-ring gland design. Dimensions and tolerances are in mm.
Chapter 2. Modeling and design
2.9 Analysis of an Ejector Pump Based on the Current Work
This section presents an analysis of a MEMS ejector pump driven by the current
device. The objective of this analysis is to derive the ejector pump specifications that were
listed in Table 1.1 and used as performance metrics for comparison against other MEMS
pumps in the literature. For the motive fluid, the proposed ejector uses a mixture of steam
and oxygen gas that is generated from the decomposition of hydrogen peroxide. The
analysis also calculates the required mass flow rate of motive fluid to provide the desired
pump performance. This required mass flow rate is then used to design the current device
as explained in the previous sections. A coupled design approach is used to determine the
pump specifications and the flow rate across the steam generator simultaneously.
Since macroscale ejector pumps typically offer pressure ratios on the order of 1:10
per stage [24], the same ratio is used here. Higher pressure ratios can be achieved by using
multi-stage pumps. This analysis determines both the mass flow rate of suction fluid that
can be pumped across a pressure rise of 1:10 and the required mass flow rate of motive
fluid. The motive fluid is taken to be a steam-oxygen mixture produced by the
decomposition of an aqueous hydrogen peroxide mixture having an overall initial peroxide
mass fraction of 83%.
Figure 2.20 is a schematic of the pump. In this configuration, the pump is used to
create a low pressure of 0.1 atm in a chamber by discharging a suction fluid from that
chamber to the atmosphere. To achieve this effect, the steam generator's effluent is used as
the motive fluid. This motive fluid is first expanded from section 0, which is the exit section
of the current device, to section 1 where the pressure matches that of the suction chamber.
.........
..
...............
......
...........................................
..
...........................................................
....
.........
.........
....
.. .......
....
.....
.
Chapter 2. Modeling and design
The motive fluid at section 1 is mixed with the low pressure suction fluid at section 2. The
pressure at both sections 1 and 2 is 0.1 atm. The mixing produces a combined supersonic
flow at section 3. The mixed flow is passed through a converging-diverging nozzle, and it
exits at section 4 where the pressure is atmospheric.
Section 0
Section 1
Section 3
Section 4
Motive fluid
w0
w11 Control
volume
Suction fluid from chamber
to be pumped down
Discharge to
I
w4
atmosphere
W2
Section 2
Figure 2.20. Schematic for ejector pump analysis
To solve this open-ended problem of determining the suction and motive fluids'
mass flow rates that would allow the realization of the above pressure ratio, the following
algorithm is used. First, the depth of all sections is fixed at 3 mm to simplify future
fabrication of the pump. Notice that this is also the depth of the reactor and nozzle of the
current device. The width of section 0 (wo) is that of the exit section of the present steam
generator's nozzle. The width of section 2 (w2) is fixed at a value comparable to that of
section 1 (wi), which is determined as described below. The width of section 4 (w4) is
matched to that of section 3 (w3), which is equal to the sum of wi and w2. Different
73
Chapter 2. Modeling and design
combinations of motive and suction fluid mass flow rates are then tried, in order to find a
combination that leads to atmospheric pressure at section 4. Iteration is required here
between this analysis and the steam generator models presented earlier, because the
motive fluid mass flow rate is needed to model and design the steam generator. The steam
generator models are used to determine the conditions at its exit. Those conditions, which
correspond to section 0 in the above figure, are needed to analyze the pump operation.
To study the above system, some assumptions are made to simplify the modeling.
Following the mixing analysis in [51], a one-dimensional compressible flow model is used.
The model is described briefly below; more detailed derivations can be found in [51].
Constant specific heats are assumed. The system walls are taken as adiabatic, and wall
shear stresses are neglected. The flow is therefore assumed isentropic everywhere except
in the mixing region, which is denoted by the dashed rectangle (control volume) in Figure
2.20. The steam generator is considered to be operating at the design conditions. The
suction fluid is assumed to be stagnant air at 250C.
The conditions at section 0 are known from the reactor and nozzle models. The
conditions at section 1 are determined as follows. The stagnation pressure is that at section
0, due to the isentropic flow assumption. The static pressure is 0.1 atm. The Mach number
M2 can be determined by using the static-stagnation pressure relation (2.34). The area Al
can then be found by applying the isentropic M-A relation between sections 0 and 1:
Chapter 2. Modeling and design
M0 A_
1 +Y
1 02
M1 A1
yo+'J
m2M1M
0M
(1±1
1+
(2.44)
n+1((2.44
2
1
In (2.44), the specific heat ratio (y) has the same value at sections 0 and 1. This ratio and the
other physical properties (specific heat and gas constant) of the flow at those sections are
determined as explained in the nozzle model. Once A2 is determined, wi can be found by
dividing Al by the depth.
The conservation of mass, energy, and momentum are then applied to the control
volume in Figure 2.20 to determine the conditions at section 3. Conservation of mass gives:
th3
(2.45)
=rh1h +rhz2,
where inis the mass flow rate at each section. Conservation of energy implies that
rjc
Tt3 =
p1Tt1 +rhzic pzTtz
.fCPp
,
(2.46)
where Te and cp are the stagnation temperature and specific heat, respectively, of the fluids
at each section. The physical properties of air at section 2 are determined using property
tables [35]. The specific heat and the gas constant (which is needed later to calculate the
specific heat ratio) at section 3 are found by mass-averaging the respective properties of
the motive fluid at section 1 and the suction fluid at section 2. The stagnation temperatures
at sections 1 and 2 and all the mass flow rates in (2.46) are known. This allows the
evaluation of Tt3.
The conservation of momentum implies that
Chapter 2. Modeling and design
P3 A3 +r 3v 3 = (P1A1 +rhiV1 )+(P2A2 +r 2 v 2 ),
(2.47)
where P,A, and V are, respectively, the static pressure, area, and velocity at each section. By
using some algebraic manipulation following [51] and realizing that the flow at section 2 is
approximately stagnant, (2.47) can be expressed as:
rh3 cp3 y3 -1
t3
h
1+yM2
-1+P2A2-
c
Ci
1(y -1)
1+ (yyM
yM 3
1+yM 2
i
1 (y-1)M
1
2
L_
2
(2.48)
i
The above expression can be used to evaluate M3. With M3 , rn3 , and Tt3 all known, the
stagnation pressure Pt3 can be determined by applying (2.39) at section 3. Note that A3 is
the sum of the areas Al and A2 .
To determine the conditions at section 4, the isentropic M-A relationship is applied
again between sections 3 and 4 to calculate M4:
j
1+
M3 A3
MA
23
1(Y3+1)
1 m2231
+
M4A4
M
(2.49)
1 r4+1)
1
1 + Y42 1 M2) 4
The stagnation pressure at section 4 is the same as that at section 3 due to the isentropic
assumption. With the Mach number and the stagnation pressure both known, (2.34) can
then be used to find the static pressure at section 4. This pressure is then compared with
atmospheric pressure. A combination of motive and suction mass flow rates that produces
atmospheric static pressure at section 4 is given in Table 2.3. The table also lists the widths
of the different pump sections.
Chapter 2. Modeling and design
To calculate the pumping capacity per unit volume, the pump's external volume has
to be estimated. The overall width and depth of the pump are taken to be the same as those
of the current device, due to the similarity in the internal lateral dimensions between the
steam generator and the proposed pump. The entire pump, however, must be longer than
the current device, since the flow from the device has to be expanded before mixing with
the suction fluid, and then the combined flow has to be passed through another nozzle. The
exact overall length depends on the detailed design of the pump, which is not covered here.
An approximate value of twice the external length of the steam generator is used. If the
volume required for 0-ring installation is included, the overall pump volume is about 3.2
cm3, which results in a pumping capacity per unit volume equal to 10-2 g/s/cm 3, as given in
Table 1.1. The volume for 0-ring installation is necessary for packaging, but can be reduced
by stacking multiple pumps. If this volume is not included in the above calculation, the
pumping capacity per unit volume increases to 1.33x10-2 g/s/cm 3. Ifthe package volume is
used as a reference instead of the pump volume, the capacity per unit volume will decrease.
Notice, however, that a package-to-pump volume increase of at least two orders of
magnitude is needed to produce a lower pumping capacity per unit volume than the state
of the art. Such an increase is unlikely with a properly-designed package. Moreover, the
ratio of package to overall pump volume can also be reduced by stacking multiple pumps in
one large package.
Chapter 2. Modeling and design
Table 2.3. Ejector pump parameters
Parameter
Value
Mass flow rate of suction fluid (kg/s)
3.2x10-s
Mass flow rate of motive fluid (kg/s)
17.5x10-5
WO (mm)
0.3
Wi (mm)
0.775
W2 (mm)
1
W3 (mm)
1.775
W4 (mm)
1.775
CHAPTER THREE
3 FABRICATION
3.1 Overview
Microfabrication was carried out at the Microsystems Technology Laboratories
(MTL) at MIT. A schematic section-view of a fabricated device is shown in Figure 3.1. The
device consists of 5 layers. Layers 1 through 4 are made from double-side-polished (DSP)
silicon wafers (Ultrasil Corp.), and Layer 5 is a capping layer made from a Pyrex wafer
(Bullen Inc.) to provide optical access to the device during experiments. The wafers for
Layers 1 and 5 are about 0.5 mm thick each, whereas those for the remaining layers are
about 1 mm thick each. All wafers are 6" in diameter. Each of the silicon wafers was
purchased with a 0.5 tm layer of oxide thermally grown on each side to protect the
surfaces during processing in order to allow successful bonding later.
Layer 1 contains the inlet holes for the peroxide and catalyst streams along with
thermocouple insertion holes to measure the silicon wall temperature. Layer 3 contains
through holes and part of the reactor and nozzle depths. Layers 2 and 4 are identical, and
contain the 100 m deep mixers on both sides of each layer in addition to the remaining
deep-etched features.
The
silicon
wafers
were
processed
via
alternating
steps
of cleaning,
photolithography, etching (wet and/or dry), and in some cases, the deposition and
patterning of additional oxide layers as masks. These wafers were then bonded together
using fusion bonding, and the silicon stack was finally bonded to the capping Pyrex wafer
via anodic bonding. Individual devices were obtained by diesawing. The entire fabrication
79
iwmw-
-- "WffMr"
-- -
- -
-
Chapter 3. Fabrication
process uses 4 masks in 16 lithography steps. The following sections describe the
fabrication process of each layer and the bonding steps. A detailed process flow is provided
in Appendix A.
LAYER 5
Peroxide
Inlet Port
Catalyst
Inlet Port
Through holes and mixers
Mixers
(four)
Reactor
Nozzle
Reactor and nozzle
Figure 3.1. Schematic section-view of a microfabricated device
3.2 Wafer-Level Die Layout
The processed wafer stack contains 28 dies, of which 14 have the "nominal" design
described in Chapter 2. Out of the remaining 14 dies, 8 have slightly modified designs to
allow the identification of any faulty section(s) in the event that the nominal devices failed
to operate successfully, and 6 are test dies designed to characterize the pressure drop and
mixing quality of the mixer sections only.
Chapter 3. Fabrication
The eight full dies with slightly modified designs are distributed as follows. (Only the
features different from the nominal design are listed).
* Two dies with shorter reactors (10 mm long), in case the nominal reactor design was
too long, thus causing significant heat losses.
* Two dies with longer reactors (20 mm long), in case the nominal reactor design was
too short, thus leading to insufficient residence time for reaction.
e Two dies with mixers having straight instead of zigzag channels, in case the zigzag
channels produced a very large pressure drop.
e
Two dies with mixers having 6 instead of 12 wall protrusions, in case the 12
protrusions produced a very large pressure drop.
The six mixer-testing test dies have no reactor or nozzle sections. They only contain the
inlet ports, the mixers, and an extra outlet port. These dies are distributed as follows:
* Two test dies with the nominal mixer design described in Section 2.3.3.
" Two test dies with straight instead of zigzag mixer channels.
" Two test dies with 6 instead of 12 protrusions on the mixer walls.
To enable the use of the test dies, a package with an extra port was also manufactured in
addition to the standard package shown in Figure 2.18. The package variations are
described in Section 4.3.
Chapter 3. Fabrication
3.3 Alignment Marks
Since the device consists of many bonded wafers, it was necessary to transfer
alignment marks to the silicon wafers before patterning the functional features. The masks
used in the subsequent processing steps have complementary features that were aligned to
these marks. A summary of the process flow for transferring these marks is shown in
Figure 3.2. The wafers were first subjected to photolithography (resist coating, prebaking,
exposure, developing, and post-baking) using thin photoresist and the mask "Alignment
Marks" shown in Figure 3.3, with a die-level zoom shown in Figure 3.4. The oxide layer was
then dry-etched, followed by etching of the silicon to a depth of 0.25 [tm. These steps were
performed on both sides of every silicon wafer. Finally the wafers were piranha-cleaned.
The mask in Figure 3.3 contains two sets of alignment marks: one set is for frontback alignment when processing wafers that are patterned on both sides, and the other set
is for aligning the different wafers during the fusion bonding step. Zooms of the alignment
marks are shown in Figure 3.5 . The mask also contains stripes that are used to guide the
diesawing process, with labels (e.g. "1"in Figure 3.4) to distinguish the different dies
described in Section 3.2.
Chapter 3. Fabrication
1. Start with a DSP silicon
wafer with 0.5 pm oxide on
each side.
2. Perform photolithography
on both sides using the mask
"Alignment Marks."
3. Dry-etch the oxide on both
sides.
4. Dry-etch the silicon on both
sides to a depth of 0.25 jim,
and clean the wafer.
*
Silicon
*
Silicon oxide
Photoresist
Figure 3.2. Process flow summary for transferring the alignment marks
............
....
. .....
..
. ...
.....
......
Chapter 3. Fabrication
Figure 3.3. "Alignment Marks" mask
Figure 3.4. Rotated die-level zoom of the mask "Alignment Marks"
....
..
..
....
..
..
.........................
..
.....
..........
.......
..
..........
Chapter 3. Fabrication
Wafer bonding
alignment mark
. . . . . . ................................................
.
Top-bottom
alignment mark
Figure 3.5. Zooms of the wafer-bonding and the top-bottom alignment marks
3.4 Layer 1
A summary of the process flow for patterning Layer 1 is shown in Figure 3.6. The
top side of the wafer was first subjected to photolithography using thick photoresist and
the mask "Holes" shown in Figure 3.7, with a die-level zoom shown in Figure 3.8. Resist was
85
Chapter 3. Fabrication
also coated on the bottom side for protection. The top oxide layer was then dry-etched to a
depth of 0.5 [rm until it was cleared from the exposed areas. The wafer was then mounted
on a quartz substrate and deep-reactive-ion-etched through the entire wafer thickness,
followed by piranha cleaning.
1. Starting with a 0.5 mm thick
with
patterned
wafer
alignment marks, perform
photolithography on the top
side using the mask "Holes."
2. Dry-etch the oxide on the
top side.
3. DRIE the silicon through the
entire wafer depth and clean
the wafer.
Silicon
Silicon oxide
Figure 3.6. Process flow summary for Layer 1
Photoresist
.......................
. ........
... ......
....
....
- . .....- -...
..........
. .........
U::r
....................
.......................
Chapter 3. Fabrication
.. ..
.......................
.................... .
.
Figure 3.7. "Holes" mask with insert showing the complementary top-bottom alignment
features
........... ...........
- - - --. .................
...
.......
Chapter 3. Fabrication
Figure 3.8. Rotated die-level zoom of the mask "Holes"
3.5 Layer 3
A summary of the process flow for patterning Layer 3 is shown in Figure 3.9. This
layer's fabrication is very similar to that of Layer 1 with one major difference. Since the
wafer in Layer 3 is twice the thickness of that in Layer 1, one coat of photoresist is not
enough to protect the thicker wafer during the silicon through-etching. Therefore, an extra
layer of 4 im thick oxide was deposited on both sides of the wafer prior to coating with
photoresist. The top oxide layer was used as a hard mask during etching, and the bottom
layer was deposited to minimize the wafer bow that could prevent successful bonding later.
The other steps are very similar to those used to pattern Layer 1.
Photolithography was carried out using the mask "Deep Features" shown in Figure
3.10, with a die-level zoom shown in Figure 3.11. Notice that in this mask, the reactor is
defined by etching its outline in a "halo-shaped" fashion instead of etching the entire
reactor area. At the end of the etch, the outlined piece falls out, clearing out the full reactor
area. The width of the halo is about 100 pim, which is closer in size than the full reactor
length and width are to most of the other features in the mask (and in the other mask used
for Layer 4). This was done on purpose to ensure etch uniformity, since DRIE causes
features with very different in-plane dimensions to be etched at different rates.
...........................................
..........
..... ......
...
.................
Chapter 3. Fabrication
1. Starting with a 1 mm thick
with
patterned
wafer
alignment marks, deposit 4
pm thick oxide on both sides.
2. Perform photolithography
on the top side using the mask
"Deep Features."
3. Dry-etch the oxide on the
top side.
4. DRIE the silicon through the
entire wafer depth and clean
the wafer.
~kY
*
Silicon
*
Silicon oxide
Figure 3.9. Process flow summary for Layer 3
Photoresist
..
........................................................................................................
Chapter 3. Fabrication
Figure 3.10. "Deep Features" mask
Figure 3.11. Rotated die-level zoom of the mask "Deep Features"
Chapter 3. Fabrication
3.6 Layers 2 and 4
Layers 2 and 4 are identical, and a summary of their fabrication process flow is
given in Figure 3.12. These layers have two sets of features of different depths: the 1 mm
deep reactor, nozzle, and hole features that extend through the entire wafer depth, and the
100 prm deep mixer features on both sides of each wafer. The top-bottom symmetry of
each of Layers 2 and 4 and the symmetry of these two layers about Layer 3 were designed
on purpose to facilitate the fabrication process and minimize the number of masks
required. Even though the mixers are relatively shallow compared to the remaining
features, they are not shallow enough to be completely and conformally covered by
photoresist if they were patterned and etched in the silicon first. This difficulty
necessitated a "nested mask" approach that used a combination of photoresist and oxide
layers as masks. By using two masking materials, it was possible to complete all of the
patterning of these masking layers before etching any features in the silicon. The complete
processing is described below.
Fabrication started by depositing 2 ptm of oxide on the top and bottom sides of each
wafer. Both sides were then coated with thick photoresist and subjected to
photolithography using the mask "All Features" shown in Figure 3.13, with a die-level zoom
shown in Figure 3.14. The oxide layers on both sides were then dry-etched, and the wafers
were piranha-cleaned. Following this step, the top surface was coated again with thick
photoresist, which conformally covered the shallow oxide layer. Photolithography was then
carried out using the mask "Deep features" that was shown in Figure 3.10 and that was
used to pattern Layer 3. The bottom silicon surface was then etched to a depth of 100 pm,
91
Chapter 3. Fabrication
producing the shallow mixers on that side. The wafer was subsequently mounted on a
quartz substrate, and the top silicon surface was etched to a depth of 800 pm. Finally, the
wafer was piranha-cleaned, and the silicon was etched an extra 100 pm from the top, which
simultaneously produced the shallow top mixers and completed the deep features.
1. Starting with a 0.5 mm thick
wafer
patterned
with
alignment marks, deposit 2
pm thick oxide on both sides.
2. Perform photolithography
on both sides using the mask
"All Features."
3. Dry-etch the oxide on both
sides and clean the wafer.
4. Perform photolithography
on the top side using the mask
"Deep Features."
Silicon
Silicon oxide
*
Photoresist
Figure 3.12. Process flow summary for Layers 2 and 4 (continued on next page)
.....
....
..
Chapter 3. Fabrication
5. DRIE the silicon on the
bottom side to a depth of 100
pm.
6. DRIE the silicon on the top
side to a depth of 800 pm.
7. Clean the wafer, exposing
the mixer features on the top
side.
I
M Silicon
8. DRIE the silicon an extra 100
pm from the top side.
*
Silicon oxide
Photoresist
Figure 3.12 (continued). Process flow summary for Layers 2 and 4
Chapter 3. Fabrication
Figure 3.13. "All Features" mask
Diesawing
plane for
this end of
the device
Figure 3.14. Rotated die-level zoom of the mask "All Features"
Chapter 3. Fabrication
3.7 Wafer Bonding
After individual processing, the silicon wafers were fusion-bonded together as
shown schematically in Figure 3.15. The silicon stack was then anodically bonded to the
Pyrex wafer as shown in Figure 3.16. To allow successful bonding, extra care was taken to
ensure wafer surface cleanliness. The silicon wafers were purchased with a thin layer of
thermally-grown oxide on both sides to protect the surfaces during processing. Once all the
individual wafers were patterned, they were piranha-cleaned and then immersed in 49%
HF acid for 5 min to remove the oxide. Immediately following this step, the wafers were
RCA-cleaned, then rinsed and dried, and finally fusion bonded. The times taken to transport
the wafers from the HF bath to the RCA cleaner, from the cleaner to the spin dryer, and
from the dryer to the bonding station were kept as low as possible (less than half a minute
each) to minimize the deposition of particulates on the wafer surfaces. Particulates could
potentially prevent successful bonding. A more detailed list of some of the practices that
were followed to ensure successful bonding is given in Appendix B. The silicon stack was
then pressed overnight and subsequently annealed at 9500C for an hour. Toward the end of
the annealing step, the Pyrex wafer was piranha-cleaned. Then the silicon stack was taken
out of the annealing tube furnace and anodically-bonded to the Pyrex immediately, and the
whole stack was again pressed for a few hours. The stack was finally diesawed, with the
device ends near the nozzles being cut along the planes shown in Figure 3.14.
.......
.. ........
-- - - --
Chapter 3. Fabrication
3. Fusion bonding of
Layer 4 to Stack 1-2-3
2. Fusion bonding of
Layer 3 to Stack 1-2
1. Fusion bonding of
Layer 2 to Layer 1
M Silicon
Figure 3.15. Schematic of the fusion bonding process
M
-
.. ..............
- -
. ..
.
- ...............................................................................................................
..............
.....
.
.....
Chapter 3. Fabrication
Layer 5
Layer 5
Anodic bonding of
Stack 1-2-3-4 to
Layer 5
M Silicon
I
Pyrex
Figure 3.16. Schematic of the anodic bonding step
Chapter 3. Fabrication
Figure 3.17 is a photograph of a microfabricated device with the nominal design.
Figure 3.18 is a photograph of a mixer-testing device. Figure 3.19 is a scanning-electronmicroscope (SEM) image showing a cross- section of the mixers and a wall protrusion. The
apparent surface irregularities in Figure 3.19 are a result of diesawing, which was needed
to obtain the cross-sectional image.
Figure 3.17. Photograph of a microfabricated device with the nominal design
10 mm
Figure 3.18. Photograph of a microfabricated mixer-testing device having an extra outlet
port in the bottom for fluid discharge, which in the fully-functional devices occurs through
the nozzle
Chapter 3. Fabrication
Wall
protrusion
Figure 3.19. SEM image of the mixers in cross-sectional view, with an insert magnifying
one mixer and showing a wall protrusion
CHAPTER 4
4 TEST RIG SETUP
4.1 Overview
To test the microfabricated devices, a rig was constructed and used to conduct the
experiments. Some of the tasks involved in setting up the test rig were designing and
machining the package in which the device sat during the experiments, implementing a
method for supplying controlled amounts of the peroxide and catalyst streams, and
interfacing the supply mechanism to the package using fittings, seals, valves, and piping. A
large part of the effort that went into all these tasks was to ensure safe operation due to the
oxidizing nature of high-test hydrogen peroxide, its strong reactivity upon contamination
by a wide range of materials, and its potential to release lots of heat during decomposition.
4.2 Hydrogen Peroxide and Safety
High-test hydrogen peroxide is a strong oxidant. For example, if 1 L of 90% HTP at
room temperature decomposes in an adiabatic closed system, it will produce 5000 L of gas
(oxygen and steam) at 7400C [69]. The high temperatures generated pose fire hazards, and
the large increase in volume can lead to explosions and detonations.
This danger is
intensified by the fact that many common materials contain significant amounts of catalysts
that can cause the decomposition of peroxide upon contact. Some examples of these
contaminant substances include wood flooring, rags, and standard clothing (cotton, wool,
leather, etc...), in addition to many commercial materials such as copper, brass, iron,
bronze, lubricating oil, nickel, gold, etc.... Peroxide is inert only to a limited number of
100
Chapter 4. Test rig setup
materials, including clean glass, passivated stainless steel, Teflon and its derivatives, PEEK,
and some Markez and Viton rubber grades.
High-test peroxide also poses significant health hazards because of its reactivity.
Skin exposure can cause blistering and burns. Eye exposure can cause corneal ulceration
and can lead to blindness. Inhalation may cause irritation to and inflammation of the nose
and throat. Accidental swallowing can lead to corrosion and internal bleeding of the
gastrointenstinal tract which could be life-threating.
The above dangers necessitate following certain practices when storing or handling
hydrogen peroxide to prevent confining, contaminating, or contacting it. A significant part
of the task of setting up the test rig consisted of planning and executing a number of safety
provisions in consultation with the peroxide vendor (FMC chemicals) to ensure safe
experimentation. These are described below.
Before purchasing the peroxide, a safety representative from the vendor company
was invited to inspect the lab and confirm its compliance with peroxide-specific safety
measures. Part of this procedure consisted of the representative giving a safety
presentation to all lab users about the possible hazards of using peroxide and ways to
counter them.
The peroxide was then purchased in a special vented 1 gallon glass bottle. A
separate vented cabinet with a lock, shown in Figure 4.1, was used for storing this bottle,
which was placed inside a 4 gallon stainless steel secondary container. Unused peroxide
was never added back to the glass bottle at the end of experiments, but was instead diluted
to about 3%concentration by adding DI water and then disposed of in the lab sink.
101
......
......
..
.......
....
. ......
.
.
........
.
Chapter 4. Test rig setup
Storage
cabinet
Vented-cap glass
bottle containing
peroxide
I:
Vents
Secondary
stainless steel
container
Figure 4.1. Peroxide storage in lab
Personal protective gear was worn during any operation involving peroxide transfer
or usage. This gear consisted of safety goggles, a polycarbonate full-face mask, and
neoprene-based overalls, apron, gloves, and boots. The safety gear is shown in Figure 4.2.
Figure 4.2. Safety gear
The experiments were carried out inside a fume hood to prevent exposure to
peroxide fumes. The testing setup was placed in a large stainless steel tray inside the hood
102
Chapter 4. Test rig setup
to capture any spillage. Since FMC Chemicals recommends using water to extinguish any
fires resulting from peroxide contamination or decomposition, a water hose was installed
and attached to the fume hood. The hood was also close to an eyewash and safety shower
station for rinsing in the event of eye or body exposure to peroxide. Blast shields made
from thick polycarbonate were attached to the hood's sash during the initial experiments.
These were disassembled later due to the difficulty of accessing the fume hood (to reach
the test rig components that needed manual adjustment, such as the plug valves described
later). To facilitate the use of blast shields in future tests, it is suggested to automate the
experiments and run them remotely, as described in Section 6.2.
The test rig components that came into contact with peroxide were all chosen to be
made of compatible materials, and were cleaned, passivated, and conditioned prior to
usage. The system was always thoroughly flushed with DI water at the beginning and end
of each experimental run. Pressure relief valves were installed to prevent pressure buildup
due to any unintended peroxide decomposition outside the device during the experiments.
4.3 Package
After an extensive materials search, Rescor 914 from Cotronics Corp. was selected
as the package material. This is a dense and vacuum-tight glass ceramic composite that is
readily machinable using standard cutting tools. This material has a low thermal
conductivity, tight machining tolerance, and a high melting point, as shown in Table 4.1
which compares the required package properties (from Section 2.8) with those of Rescor
914. In addition, this material is compatible with peroxide. This fact was demonstrated
experimentally by soaking a sample of Rescor 914 for 3-4 hours in 30% peroxide and
103
Chapter 4. Test rig setup
observing very little reaction. The logic behind using 30% concentration was that this
concentration was high enough to provide a visual indication of incompatibility in the form
of bubbles. At the same time, 30% concentration was low enough to ensure safety, by
producing much lower reaction temperatures and gas volumes than 90% concentration in
the event of decomposition. All these characteristics made Rescor 914 suitable as a package
material.
Table 4.1. Comparison of the required package properties and those of Rescor 914
Property
Thermal conductivity (W/m-K)
Required value
(from Section 2.8)
s 0.8 - 3.2
Melting point (OC)
400
Tolerance (mm)
s 0.05
Value for Rescor 914
0.4
-500
~0.05
The package was manufactured at Ferro-Ceramic Grinding Inc., which specializes in
the machining of ceramics. The actual package is shown in Figure 4.3. Two views of each of
the top and bottom parts are shown. The "inner" surface of each part is the surface that
comes in contact with the device. The thermal resistance of the whole package was found
to be on the order of 8 K/W using (2.43). Zirconia screws and nuts (Ceramco Inc) with low
thermal conductivity were first used to clamp the two parts of the package together with a
minimal drop in the package's thermal resistance due to clamping. These screws, however,
were found to be too brittle to allow tight fastening of the package, and they were later
replaced with stainless steel screws (McMaster-Carr). The package-to-device interface
104
Chapter 4. Test rig setup
consisted of 0-rings made from Markez Z1030 (Marco Rubber). This material is compatible
with both the high-test peroxide and the ferrous chloride catalyst and can sustain the
predicted device temperatures.
Top part - inner surface
Bottom part - inner surface
-
Top part - outer surface
Bottom part - outer surface
10mm
Figure 4.3. Machined package made of Rescor 914
In addition to the above "standard" package design, two modified designs for the
bottom part were also manufactured from the same material and used with the standard
top part. One of the modified bottom designs was machined after finding out that
thermocouple insertion through the original bottom half of the package was very
challenging. In the original design, the device and the package are first assembled, and then
the thermocouple wire is inserted. The thermocouple must first pass through the 12.7 mm
105
Chapter 4. Test rig setup
deep holes in the package before it reaches the device surface. Then, the thermocouple has
to be inserted through the narrow device holes, whose diameters are designed to match
those of the thermocouple wires to achieve an interference fit. This task was found to be
very cumbersome and time-consuming. The new design, shown in Figure 4.4, facilitated
thermocouple use and insertion by having long slits along its interior surface instead of
round through holes terminating at the device's surface. With the new design, the
thermocouples are inserted into the device, bent, and then passed through the slits before
clamping the package. This makes the insertion process easier. Figure 4.5 is a schematic
explaining the thermocouple insertion methods for the original and modified designs.
Figure 4.4. Modified package bottom half for facilitating thermocouple insertion
106
.:
....
........
......
....
.......
.
. ................
I..............................................
I...............
Chapter 4. Test rig setup
(a) Original package design. In
this
design,
the
thermocouple is attached
to
the device
after
assembling the package.
The thermocouple must
pass through the 12.7 mm
deep holes in the package
before it can be inserted
into the narrow (0.5 mm
diameter) device holes.
Narrow
device hole
Deep package holes
Bent thermocouple,
(b) Modified package design
for easier thermocouple
insertion. In this design,
the
thermocouple
is
attached to the device
before assembling the
package.
The
thermocouple is inserted
into the device holes,
bent, and then passed
through the slits along the
package surface before
clamping the package.
Figure 4.5. Schematic explaining the difference in the thermocouple insertion methods
between the original and modified designs of the bottom half of the package
The second modified design of the bottom half of the package, shown in Figure 4.6,
allowed the usage of the mixer-testing devices by having an extra outlet port. This design
107
..
....
...................
::::
...............
Chapter 4. Test rig setup
had no thermocouple insertion features. Note that in both Figures 4.4 and 4.6, the red
discoloration of the package was due to the settling of the post-catalysis ferrous chloride
residue and its possible reaction with the package material.
Figure 4.6. Modified package bottom half for usage with the mixer-testing devices
4.4 Test Rig Components
The main components of the test rig are shown schematically in Figure 4.7. Syringe
pumps (Chemyx Inc.) were used to control the mass flow rates of the peroxide and catalyst
solutions. The syringes are made of SS316 stainless steel and have Viton 0-rings. The
volumes of the peroxide and catalyst syringes are 100 mL and 20 mL respectively. The
plunger force capacity and syringe dimensions were chosen so that the pressures produced
inside the syringes were almost the same. Pressure relief and plug valves (Swagelok) with
stainless steel bodies and Viton 0-rings, in addition to PEEK check valves (Kinesis Inc),
were installed in the flow path of each of the peroxide and catalyst streams. The pressure
relief valves were used to prevent peroxide confinement and pressure buildup inside the
system in the case of accidental decomposition in the flow paths leading to the package.
These relief valves were calibrated to open at ~ 10 atm. The plug valves are manual on/off
valves that were used to either connect the streams to the package or to purge each stream
108
Chapter 4. Test rig setup
into a separate waste container during shutdown. The check valves were installed very
close to the package and were used to prevent either stream from flowing into the piping of
the other stream through the package, especially during startup. Stainless steel tees and
adapters (Swagelok), PEEK adapters and fittings with stainless steel ferrules (Kinesis Inc),
and semi-rigid Teflon tubing (3.2 mm OD, 1.5 mm ID, from Kinesis Inc.) were used to
connect these components to one another, to the syringe pumps, and to the package. Figure
4.8 is a photograph of the clamped package connected to the Teflon tubing via the PEEK
fittings. All components were cleaned and passivated before exposure to peroxide, as
described in Section 4.5.
Large glass beakers were used to collect the peroxide, catalyst, and reaction product
wastes, which were then disposed of following the EHS (Environment, Health, and Safety)
department's waste management procedures. The syringe pumps, plug and pressure relief
valves, and peroxide and catalyst waste collectors sat inside a large stainless steel tray to
capture spillage and prevent contamination of the rest of the fumehood. Inside this large
tray, the syringe pumps were placed on another inverted perforated stainless steel tray to
keep them from getting wet if the large tray needed to be flushed with water. The package
was raised above the rest of the system and was held in place by a clamping mechanism.
The post-decomposition waste collector sat under the package. The tubing was attached to
the package from below, and the device's Pyrex layer faced upwards. An optical diagnostic
system consisting of a camcorder, lens, and illumination source was used to look through
the Pyrex and record videos of the device during operation. Further characterization was
obtained using temperature measurements and refractometry as described in detail in
109
Chapter 4. Test rig setup
Section 5.3. Figure 4.9 is a test rig photograph taken with the setup placed outside the
fumehood to obtain a good picture quality. The beakers for collecting the peroxide, catalyst,
and product waste are also not shown. During the peroxide experiments, the setup was
inside the hood at all times, and the waste collectors were always used.
PRV pressure relief valve
PV Plug (on/off) valve
waste
PRV
CV: Check valve
High-test
Peroxide
p
PV
CV
Package
PRV
fProduct
waste
collector
Ferrous
chloride
lI~ldt
tetrahydra
aIL
te
PV
Figure 4.7. Schematic of test rig setup
110
.
Chapter 4. Test rig setup
C
Teflon
Package
tubing
Screws
PEEK
fittings
Figure 4.8. Assembled package, fittings, and tubing
Syringe pumps
Lens system
a
Large containment
tray
Inverted
perforated tray
Camcorder
N
Pressure relief and
plug valves
Package and check
valves
Figure 4.9. Photograph of the test rig highlighting the main components
111
..
.............
Chapter 4. Test rig setup
4.5 Component Passivation
Passivating the components that will handle high-test hydrogen peroxide is the
process of removing contaminants (such as iron or other atoms) from the components'
surfaces. This was carried out to reduce the surface activity of the components and to
increase their chemical stability upon contact with the peroxide. The process consisted of
cleaning the components, then chemically dissolving any contaminants by acid immersion,
and finally conditioning the parts with 30% peroxide to form a protective layer. The
syringes, valves, connectors, and fittings were all passivated prior to performing any
experiments using this procedure [70], which is described in detail below.
All components were disassembled into their individual (usually single-piece)
constituent parts. The parts were then divided into two large glass beakers, one containing
all the metal parts, and the other containing all the nonmetal (thermoplastic and elastomer)
parts. Both beakers were filled with high strength ultrasonic cleaner solution and placed in
an ultrasonic bath for two hours. After this cleaning process, the contents of both beakers
were rinsed thoroughly with DI water. The beaker with metal components was then filled
with 70% nitric acid (Transene Inc.), and the other beaker was filled with 35% nitric acid.
The parts were immersed in the acid solutions for 1.5 - 2 hours and then rinsed again using
DI water. The contents of both beakers were then conditioned with 30% hydrogen
peroxide (Transene Inc.) for 12-16 hours, while watching closely for decomposition, which
was usually observed as bubble formation. After all decomposition ceased, the parts were
cleaned one last time with DI water and reassembled while wearing gloves to prevent
contamination. When lubrication was necessary, a special fluorolube (ChemPoint) that was
112
Chapter 4. Test rig setup
compatible with peroxide was used. The components were finally air-dried, placed in a
sealed plastic bag, and considered ready for usage with high-test peroxide. The
components were always handled using gloves from that point onwards.
113
CHAPTER FIVE
5 TESTING
5.1 Overview
After the test rig was set up, experiments were carried out to test the fabricated
devices. The purpose of these tests was two-fold: to demonstrate successful operation of
the devices, and to provide experimental verification of the multi-domain model that was
discussed in Chapter 2. The devices' performance was benchmarked using two criteria: the
degree of decomposition of the peroxide and the degree of vaporization of the water. Full
decomposition and complete vaporization were desired for successful demonstration.
Temperature and species' mass fraction measurements were taken to assess the model
quantitatively. Characterization tools used include an optical diagnostics system for visual
inspection, thermocouples for temperature measurements, and a refractometer for the
mass fraction analysis.
5.2 Experiments with Mixer-Testing Devices
Before the nominally-designed devices were tested using peroxide, experiments
using dyed water were carried out on the mixer-testing devices that have the standard
mixer design. The purpose of these experiments was to assess the capability of the setup to
handle the desired mass flow rates without leakage and also to characterize the mixing
quality. Using dyed water instead of peroxide and catalyst allowed this assessment to be
carried out quickly without having to implement all of the time-consuming safety measures
required to use peroxide.
114
Chapter 5. Testing
The mixers have comparable viscous resistance to the dyed water and peroxidecatalyst solutions due to the similarity in the hydrodynamic properties (such as density and
viscosity) of these solutions. The exit pressure in both the mixer-testing devices and the
nominally-designed devices is close to atmospheric, and the pressure drop is dominated by
the viscous losses in the mixer section in both cases. Therefore, the inlet pressure required
to supply a given mass flow rate of water to the mixer-testing device will be close to that
required to supply the desired mass flow rate of peroxide and catalyst to the regular device.
The objective here was not to quantitatively measure the supply pressure, but rather to
qualitatively confirm whether it could be handled by the various setup components
(syringes, valves, 0-rings, etc...) without leakage or failure. For this purpose, two water
streams were flowed through the two inlet ports of the mixer-testing device at different
flow rates. No syringe pump stalling and no leakage were observed in the system for total
mass flow rates up to at least 2-3 times the design value.
To test the mixing quality, yellow and blue dyed water streams were flowed through
the mixer-testing device at the design mass flow rates of the peroxide and catalyst
solutions, respectively. The different colors of the water streams allowed the assessment of
the mixing quality using visual inspection. Again, the objective was to obtain a qualitative
rather than a quantitative assessment.
The blue-yellow dyed water system and the
peroxide-catalyst system had similar hydrodynamic properties (which influence the mixing
flow regime, as described in Section 2.3) and binary diffusion coefficients. Matching of the
flow rates and physical properties of the two systems ensured accurate simulation of the
115
Chapter 5. Testing
mixing under design conditions. A magnifier was used to look through the Pyrex. A green
solution was observed at the exit of the mixer section, which confirmed successful mixing.
5.3 Experiments with Nominally-Designed Devices
5.3.1 Experimental conditions
To demonstrate successful operation, the nominally-designed devices were tested
using high-test peroxide and a catalyst solution. First, sodium permanganate was selected
as the catalyst, since its activity is reported to be the highest among peroxide catalyst
solutions [43]. The permanganate solution, however, was quickly found to clog the devices
due to the large insoluble particles that it contained. It was then replaced with ferrous
chloride tetrahydrate solution, which is reported to have a moderately high activity
without containing large insoluble particles. With the ferrous chloride, no clogging was
observed during the experiments.
The devices were tested at the design conditions, namely by flowing a 90% HTP
mixture at 7.0 mL/min and an 80% saturated ferrous chloride tetrahydrate solution at 0.5
mL/min. These conditions corresponded to a total mass flow rate of 17.5x10-5 kg/s and an
overall initial peroxide mass fraction of 83%. Apart from demonstrating successful
operation, the purpose of the experiments was also to provide comprehensive verification
of the model. To that end, two other experiments were carried out using the same mass
flow rate but different overall initial peroxide mass fractions: 74% and 71%. These two
values were chosen because they bound the range of nonlinear behavior seen in the
parametric study results (e.g. Figures 2.11, 2.12, and 2.14), and together with the design
116
Chapter 5. Testing
conditions could provide very useful insight into the model's validity. To obtain the lower
peroxide mass fractions, the 90% HTP mixture was diluted by adding DI water while
making sure that two quantities were kept constant. The first was the catalyst flow rate,
which was kept fixed so that the number of catalyst particles available for facilitating the
reaction was not altered between different experiments. The second was the total mass
flow rate, which was fixed to ensure successful sonic and supersonic acceleration of the
flow in the fixed-geometry nozzle. The input to the syringe pumps is volume rather than
mass flow rate. Since the density of the peroxide-water mixture changes with dilution, care
was taken in calculating the required volume flow rate of diluted HTP for each experiment.
The conditions for all three experiments are shown in Table 5.1.
Table 5.1. Conditions during the peroxide experiments on nominally-designed devices
HTP stream Catalyst stream
volume flow
volume flow
rate (mL/min) rate (mL/min)
Peroxide mass
fraction in HTP
stream (%)
Overall initial
peroxide mass
fraction (%)
Experiment 1
(Design
conditions)
7.0
0.5
90.0
83
Experiment 2
7.3
0.5
79.9
74
Experiment 3
7.4
0.5
76.7
71
5.3.2 Visual inspection
The device was visually inspected during operation to check for vaporization and to
qualitatively verify the model. The inspection was carried out using a camcorder, lens
117
Chapter 5. Testing
system, and illumination source to record videos in the case of experiment 1, and by eye in
experiments 2 and 3. For experiments 1 and 2, inspection showed that the device
experiences three distinct temporal stages after the system is turned on. First, a lot of liquid
is emitted from the device, in addition to some gas, as seen in Figure 5.1a. This is because
the device walls are still very cold, and consume most of the heat generated by the reaction.
As a result, only a small portion of the generated heat is left for vaporizing the liquid
species. The walls soon start heating up, and all the liquid inside the device starts getting
vaporized, leading to the second stage. The nozzle accelerates the vapor by reducing its
static pressure and temperature. During this stage, the temperature attained by the flow
due to the reaction before entering the nozzle is still not as high as desired. As a result, the
flow expansion in the nozzle produces static temperatures that are comparable to the
boiling point of water at the given static pressures, especially near the exit. This causes the
vapor to condense inside the nozzle or shortly upon exiting it. The effluent in this stage
looks "misty" due to the condensed water droplets carried by the remaining high-speed gas
mixture, as seen in Figure 5.1b. In the third and final stage, the device walls reach the
desired steady-state temperature dictated by the reaction. The flow temperature is now
high enough at the end of the reactor to allow flow expansion in the nozzle without
approaching the boiling point of water. As a result, the effluent remains completely gaseous
with no visible condensation near the exit, as seen in Figure 5.1c. Even though the effluent
is not visible in this figure, the device performance at this stage is actually the best among
all three stages, since visibility indicates condensation. In experiment 1, the first two stages
(combined) lasted for about 30s, and the third (steady-state) stage extended for the
118
Chapter 5. Testing
remainder of the syringe pump cycle (about 11.5 minutes). This visual assessment served
as a first indicator of the complete vaporization of the effluent at steady state for
experiments 1 and 2, as predicted by the model.
For experiment 3, after the initial startup stage in which liquid bubbles were
emitted, the effluent changed into the misty form mentioned above. This is due to some
condensation taking place in the nozzle, as described in the previous paragraph. For this
case, however, the effluent remained in the misty form until the end of the experiment.
This observation agreed with the model, which for the conditions of experiment 3
predicted condensation inside the nozzle at steady state.
(a)Startup stage
(b) Intermediate transient stage
Bubbles
-
Misty
effluent
4 mm
(c) Steady-state stage with no visible condensation
Figure 5.1. Frame-grabs of the device during experiment 1 showing the effluent in the
different stages: (a)startup, (b) intermediate transient period, and (c) steady state
119
Chapter 5. Testing
5.3.3 Refractive index analysis
Refractometry was used to study the extent of peroxide decomposition. The value of
the refractive index of a liquid peroxide-water mixture changes when the fractions of
peroxide and water change. The refractive index can therefore be measured to determine
the mass fraction of peroxide in the mixture. For this purpose, a hydrogen peroxide
refractometer (Atago Co., model PAL-39S) was used. The refractometer takes a small
sample of an aqueous peroxide mixture and passes a light beam through it to measure its
refractive index, which is then correlated with the peroxide mass fraction. The instrument
can be used with mixtures containing up to 50% peroxide by mass, provided that the only
other compound in these mixtures is water.
To perform the above analysis on the gaseous effluent during experiments 1 and 2,
the procedure shown in Figure 5.2 was used. The effluent was captured in a glass beaker at
a distance where all the solid catalyst had fallen off due to gravity. This distance was
determined using trial and error, by capturing the effluent and observing its color after
condensation. Any dark brown color was associated with the catalyst, so the distance from
the device to the beaker was increased until the condensate became clear. Once the correct
distance was determined, the effluent was captured in a clean cold beaker and allowed to
condense. A sample of the clear condensate was then analyzed using the refractometer to
determine the amount of peroxide present. Figure 5.3 plots the experimentally-measured
peroxide mass fractions in the effluent for experiments 1 and 2. The figure also compares
the measured values to the simulation results at the reactor exit. For complete
decomposition in the reactor, the mass fraction of peroxide is ideally zero at both the
120
Chapter 5. Testing
reactor and nozzle exits. Very good agreement is observed between the model and
experiments 1 and 2. This agreement validates the model and confirms the complete
decomposition of peroxide down to an initial mass fraction of about 74%, as predicted.
For experiment 3, it was not possible to obtain a clear condensate in the beaker. The
condensate always had a brown color, which indicated the presence of catalyst. Unlike
experiments 1 and 2 where the solid catalyst falls off from the gaseous products, some
catalyst is always dissolved in the liquid-gas effluent of experiment 3. The refractometer
readings are only meaningful when the mixtures analyzed consist exclusively of hydrogen
peroxide and water. This was not the case for experiment 3. As a result, it was not possible
to obtain an accurate measurement of the mass fraction of undecomposed peroxide for this
experiment using the above approach. Nevertheless, the refractometer was used for
qualitative model verification. A nonzero (-10%) value was measured for the peroxide
concentration. The color of the condensate provided qualitative proof of the presence of
liquid catalyst in the effluent. This confirmed the model-predicted conclusion that
condensation occurred inside the device under the conditions of experiment 3. The fact
that the refractometer measured a nonzero value is consistent with the presence of liquid
catalyst in the effluent as well as with the presence of undecomposed peroxide, as was
predicted by the model.
121
.....
.. .....
... .........
...................
.............
................
Chapter 5. Testing
1.Capture the effluent in a cold glass
beaker at a distance where all the solid
catalyst has fallen off due to gravity.
2.Cap the beaker and
let the effluent
condense.
3. Pour about 0.2 mL of
the condensate into the
refractometer seat and
analyze it.
Figure 5.2. Procedure used to perform the refracive index analysis on the effluent
Owrall initial peroxide mass fraction (%)
Figure 5.3. Plot of the peroxide mass fraction at the reactor exit (from the simulation) and
at the nozzle exit (from experiments 1 and 2) versus initial peroxide mass fraction
122
Chapter 5. Testing
5.3.4 Device wall temperature measurements
Thermocouples were used to measure the temperatures of the silicon walls of the
device during the experiments. Both grounded and exposed junction styles were tried with
the new package design that was shown in Figure 4.4. In the grounded variety, the
thermocouple wire is completely covered by a sheath to which the junction is physically
attached. This sheath, and not the junction, comes into direct contact with the medium
whose temperature is being measured. In the exposed variety, the junction protrudes
outside the tip of the sheath and contacts the medium directly. Exposed thermocouples
offer faster response and potentially more accurate measurements, but are more
susceptible to corrosion, especially in wet media. When grounded thermocouples were
used, the measured temperatures were found to be less than those predicted by the model.
This was attributed to the fact that the thermocouple sheath contacted both the device wall
and the interior package surface, and therefore equilibrated at some intermediate
temperature. Despite the discrepancy, this measurement confirmed that the wall was
almost isothermal, since temperatures measured at different locations along the device
length all agreed to within 50C. In drawing this conclusion, it is implicitly assumed that the
contact configuration was identical for the measurements that were taken at different
locations. This is a reasonable assumption, since the type and diameter of the
thermocouple wire used, the device hole geometry, the width of the slit in the package, and
the manner of insertion were all the same between the different locations. When exposed
thermocouples were used, it was not possible to make any temperature measurements
because the thermocouple junctions got attacked by the catalyst residue during startup. An
123
Chapter 5. Testing
accurate measurement was finally obtained by bringing the grounded thermocouple tips
into contact with the exposed surface of the device at the exit section. Figure 5.4 plots the
experimental results obtained using this method and compares them with the wall
temperatures predicted by the parametric study. The uncertainty in the measured values
is due to the thermocouple error limits. The measured temperatures agree with the model
predicitions to within 2%, providing further model verification. It should be noted that the
simulation curve in Figure 5.4 follows the trend of the effluent static and stagnation
temperatures explained in Section 2.6. This similarity in shape further justifies the
approximate assumption of linear proportionality between wall and effluent temperatures
that was made in Section 2.6 to justify the shapes of the effluent temperature curves.
650Simulation
600-
Experiments
+
550E
500-
450-
400I
71
I
I
I
1
I
I
|
1
I
|
72 73 74 75 76 77 78 79 80 81 82 83
Overall initial peroxide mass fraction (%)
Figure 5.4. Plot of the silicon wall temperature from the simulation and the experiments
versus initial peroxide mass fraction
124
Chapter 5. Testing
5.3.5 Effluent temperature measurements
Grounded thermocouples were also used to measure the effluent temperature at a
distance of 1-1.5 mm from the nozzle exit. The model variable against which this
measurement is compared is the "adiabatic wall temperature" Taw of the flow at the
thermocouple location [65]. This quantity describes the temperature of a heated highspeed flow when it comes to rest at a solid surface placed in its way. The temperature Taw
therefore represents the maximum temperature that the solid can attain, and it provides an
upper bound for the thermocouple reading.
For the present range of measured
temperatures, the thermocouple error specification is no larger than a few degrees, which
suggests that the difference between the reading and the flow's (adiabatic wall)
temperature should be small. The quantity Taw is given by:
1 rv2
Taw = Texp +-- _x.
2 cP
(5.1)
In (5.1), Texp and Vexp are the expanded flow's static temperature and velocity, respectively,
at the thermocouple location, r is the recovery factor, and c, is the average specific heat.
Due to viscous losses, not all of the flow's kinetic energy is converted to thermal energy
when the flow comes to rest on the solid surface. The recovery factor is a measure of the
fraction of the kinetic energy that is recovered as thermal energy. For inviscid flow, the
recovery factor is 1, and the adiabatic wall temperature becomes the stagnation
temperature at the location of interest. More generally, for turbulent gas flows such as the
effluent in experiments 1 and 2, the recovery factor is given by [65]:
125
Chapter 5. Testing
r = Pr1 /,
(5.2)
where Pr is the Prandtl number of the gas mixture. For (5.1) and (5.2) to hold, C,and Pr
must be evaluated at the reference temperature Tref given by [65]:
Tref = Tex, +0.5(Toid - Texp)+0.2 2(Tw - Texp ),
(5.3)
where Tsolid is the actual surface temperature of the solid placed in the flow. For the current
experiments, both Taw and TsoIid are expected to be very close to the temperature Tmeas
measured by the thermocouple. This allows Tref to be approximated by:
Tref = Texp +0 .7 2(Tmeas - Texp) .
(5.4)
The consistency of the assumption that Taw and Tmeas are not very different is checked after
calculating Taw.
For experiment 3, the effluent mixture is liquid-gas. Expression (5.2) is therefore
invalid for the conditions of this experiment, and the recovery factor is varied between 0
and 1 in the corresponding calculation.
The velocity Vexp is determined by applying the conservation of mass between the
nozzle exit and the thermocouple location, assuming that the bulk one-dimensional flow
approximation holds in this region. This results in the expression:
(PVA)e =(PVA)exp,
(5.5)
where the subscript e refers to the flow conditions at the nozzle exit section, and exp refers
to the expanded flow conditions at the thermocouple location. The densities at these
locations are almost identical. This can be seen from the ideal gas law. The static pressure
at both locations is ambient. The variation in the static flow temperature between the two
126
Chapter 5. Testing
locations is small. This is because the timescale of heat transfer from the flow to the nearby
air, which is ultimately dominated by conduction in the still air, is much larger than the
timescale of motion of the high-speed flow itself. The densities can therefore be dropped
from (5.5), making the velocity ratio inversely proportional to the area ratio:
Vexp
Ae
Ve
Aexp
(5.6)
The area and flow velocity at the nozzle exit can be determined from the nozzle model. The
flow velocity at the thermocouple location can therefore be determined if the flow area at
that position is known. This area is determined using the experimental findings of Quinn
[71], who studied the free expansion of a turbulent air jet exiting from a rectangular
channel. In this study, the jet half-velocity width and depth, which together provide a
measure of the change in the jet area due to the free expansion, were evaluated at different
distances from the channel exit. Plots were then generated in [71] to correlate, in
dimensionless form, the lateral dimensions of the expanded free jet to the distance from
the exit section and the channel dimensions. These plots are used in the present work after
accounting for the difference in the exit channel's aspect ratio between the above study and
the current work. The area ratio in (5.6) is found to be in the range 0.59 - 0.72. The range in
values is due to the uncertainty in the thermocouple location (between 1 and 1.5 mm away
from the nozzle exit).
Using the nozzle exit velocity given by the nozzle model, the
expanded flow velocity is then determined using (5.6).
The static flow temperature at the thermocouple location is approximated by the
model-predicted nozzle exit temperature, as explained above. With this temperature, the
127
Chapter 5. Testing
expanded jet velocity, and the recovery factor all known, (5.1) is used to calculate the
adiabatic wall temperature.
Table 5.2 compares the measured effluent temperatures with the predicted
adiabatic wall temperatures at the thermocouple location. The uncertainty in thermocouple
placement affects the expanded jet velocity and produces a range of predicted
temperatures, rather than a single predicted temperature, in all three experiments. For
experiment 3, there are two additional sources of error that contribute to this range. The
first source is the uncertainty in the recovery factor, which was varied between 0 and 1 for
the temperature calculation in experiment 3. The second source is the uncertainty in the
quality of the steam in the effluent, since for this experiment only, some condensation
occurs in the nozzle. The exact value of the steam quality is unknown, so a relatively large
range of steam qualities between 20% and 80% is used in the calculation to span a wide
spectrum of possible cases. This range leads to a variation in the properties of the effluent,
which are calculated as the mass-weighted averages of the individual properties of steam,
liquid water, and oxygen. It also leads to a variation in the effluent velocity at the nozzle
exit and at the thermocouple location. The predicted temperatures reported for experiment
3 reflect all these variations. The recovery factors are also reported in Table 5.2.
128
Chapter 5. Testing
Table 5.2. Recovery factors, predicted adiabatic wall temperatures, and measured
effluent temperatures for the three experiments
Predicted range in
Recovery
factor
Temperature
adiabatic wall
measured by
temperature at
thermocouple (K)
thermocouple location (K)
Experiment 1
0.943
623.1 - 643.4
623.0 ± 2.6
Experiment 2
0.947
484.2 - 500.6
499.0 ± 2.2
Experiment 3
0-1
353.0 - 375.8
360.0 ± 2.2
Despite the uncertainties in the temperature calculation, very good agreement is
observed between the predicted and measured temperature values. This provides further
confirmation of the model's validity.
The above analysis is also used in reverse to confirm full water vaporization by the
device under the design conditions of experiment 1. The temperature measured by the
thermocouple is now used as the adiabatic wall temperature in (5.1), and r and Vexp are
calculated following the same procedure as above to determine the experimental static
flow temperature at the nozzle location. This temperature is found to be in the range
562.9K-583.2K, which is well above the saturation temperature of water (364.8 K) at the
exit conditions. This finding implies that all the H20 exiting the nozzle is steam, which
confirms full vaporization of the liquid water inside the device under the design conditions.
129
Chapter 5. Testing
5.3.6 Summary of the implications of the experimental results
The peroxide experiments had two main objectives. The first objective was to
demonstrate successful system operation, defined by achieving complete peroxide
decomposition and full water vaporization, under design conditions. The second objective
was to provide comprehensive model verification over a range of experimental conditions.
For the first objective, the refractive index results indicated complete peroxide
decomposition,
and
the
visual
assessment
and
the
effluent
temperature
measurements/analysis confirmed full vaporization. For the second objective, the
measured peroxide mass fractions and device wall and effluent temperatures were
compared against their corresponding model-predicted values, providing quantitative
model verification. The visual assessment also provided qualitative model verification, via
the occurrence or absence of visible condensation in the device under different conditions,
as predicted by the model. Good agreement between experiment and theory confirms the
accuracy of the detailed physical modeling, which was a key enabler of successful
operation.
130
CHAPTER SIX
6 CONCLUSIONS
6.1 Summary and Important Findings
This thesis presented a MEMS device that utilizes the decomposition of hydrogen
peroxide to generate steam. The device operates by decomposing high-test hydrogen
peroxide using ferrous chloride catalyst. The decomposition produces oxygen gas and
water, and it also releases heat which subsequently vaporizes the water into steam. The
heated gaseous mixture is then accelerated through a nozzle. The resulting high-speed flow
can be used to drive MEMS ejector pumps which are capable of pumping gases at high mass
flow rates. This functionality, i.e. the high flow rate pumping of gases, is currently lacking in
MEMS technology and can have a very significant impact on many MEMS systems. Analysis
of an ejector pump based on the current work predicts that the pump capacity per unit
volume will be on the order of 10-2 g/s/cm 3, which is two orders of magnitude higher than
the state of the art. A pressure ratio of 1:10 is possible with this mass flow rate and one
stage of the pump; higher pressure ratios can be achieved by using multiple stages.
Hydrogen peroxide is a very attractive green candidate for generating steam to drive this
kind of pump. By virtue of its high density and its chemical energy, peroxide allows the
realization of compact ejector pumps that do not require external power sources.
To our knowledge, previous attempts at demonstrating similar MEMS devices that
generated steam from the decomposition of hydrogen peroxide have generally been
unsuccessful. This is attributed to a number of reasons. One is the use of heterogeneous
catalysts which are susceptible to poisoning and can lead to clogging of the devices. A
131
Chapter 6. Conclusions
second reason is improper thermal management which has usually lead to significant heat
losses from the device. These losses can compromise the device's performance, since the
retention of the heat generated due to decomposition is necessary for sustaining the
reaction and vaporizing the water produced. A third reason is the use of incomplete
physical models that failed to capture all the relevant physics, such as heat transfer and
boundary layer formation.
Successful demonstration of the current device is a result of the implementation of a
number of design features to handle the above challenges. The catalyst challenge was
addressed by using a homogeneous (liquid) catalyst solution. A liquid catalyst eliminates
the poisoning problem by being continuously supplied, and it also allows wide mixer
channels that are less susceptible to clogging. To enable successful use of this kind of
catalyst, four parallel engulfment flow mixers were designed to allow fast mixing of the
catalyst and peroxide streams. Each mixer has a residence time of about 1 ms, which
matches the mixing timescales of engulfment flow mixers in the literature that have been
demonstrated to achieve high mixing qualities. The current mixers also have zigzag
channels and wall protrusions, which have been shown to improve mixing. The thermal
management challenge was addressed by using a thermally-insulating package to minimize
the heat losses while providing piping ports. The incomplete modeling challenge was
addressed by developing a comprehensive, multi-domain model that accounts for the
different physical phenomena taking place inside the device and package. These
phenomena include chemical reaction, phase change, compressible flow, heat transfer
inside the device and to the surroundings, and boundary layer formation in the nozzle. By
132
Chapter 6. Conclusions
using these models, the device was designed to be large enough in order to prevent heat
losses from compromising the operation, which is a risk at very small sizes. Modeling also
allowed the estimation of the maximum allowable thermal conductivity of the package,
which was found to be in the range 0.8-3.2 W/m-K. The actual package material used,
which is a machinable mica-based ceramic composite, has a conductivity of 0.4 W/m-K. The
model was also used to estimate the thickness of boundary layers in the nozzle. The
formation of these layers was then compensated for by increasing the nozzle width to keep
the mass flow rate unaltered. Finally, the model was used to conduct a parametric study of
the device over a range of operating conditions. The study was later verified
experimentally. Close agreement between theory and experiments indicated that the used
model successfully captured all the relevant physics that took place during operation.
The device was microfabricated by bonding four silicon wafers and one Pyrex wafer.
The silicon wafers were first bulk-micromachined using deep reactive ion etching, and then
they were fusion-bonded. This allowed the realization of the inlet ports, the mixers, the
reactor, the nozzle, and the thermocouple insertion holes. Finally, the Pyrex wafer was
anodically bonded to the silicon stack to cap the structure and provide optical access
during experiments. A number of measures (see Appendix B) were followed to allow
successful bonding.
The test rig consisted of syringe pumps to control the peroxide and catalyst flow
rates, in addition to pressure relief, plug, and check valves. Despite being nontoxic,
hydrogen peroxide can pose some health hazards due to its strong oxidizing capabilities. As
a result, a number of features were implemented to ensure safe peroxide storage and
133
Chapter 6. Conclusions
handling. A lot of emphasis was placed on using compatible materials for the test rig
components, to prevent contamination and possible decomposition of the peroxide outside
the device.
Primary experiments were conducted to qualitatively study the mixing quality in
test devices. These devices only contained the mixer section in addition to inlet and outlet
ports. Good mixing quality was observed. Experiments on the nominally-designed fullyfunctional devices followed. The objectives of these experiments were two-fold: to
demonstrate successful operation and to comprehensively verify the model. Various
characterization tools were used, including visual inspection, refractive index analysis, and
temperature measurements. These tools indicated that full peroxide decomposition and
complete water vaporization occurred inside the device, which confirmed that the device
was operating successfully. In addition, the experimental results agreed very closely with
the model, which verified the model's validity as mentioned previously. The realization of
this steam-generating device, which other groups have unsuccessfully attempted to
demonstrate for over a decade, represents a breakthrough that can have a huge impact on
the MEMS and microfluidics fields.
6.2 Challenges and Future Improvements
Despite the successful demonstration of the device's performance, some challenges
were encountered during testing. This section describes those challenges and suggests
ways for improving future experiments.
The first challenge was the clogging of devices between experimental runs. New
devices were observed to work properly as long as the syringe pumps were operating.
134
Chapter 6. Conclusions
When the syringes ran out of peroxide or catalyst, the experiments had to be stopped
temporarily to refill those syringes. Attempting to run the experiments again with the same
device after refilling the syringes was found to set off the pressure relief valves, which
would start leaking as soon as the flow in the supply pipes reached the device. Replacing
the old device with a new one alleviated this problem. These observations indicated that
the devices were getting clogged, not during each experiment but between different
experimental runs. This is attributed to the settling of particulates inside the device at
shutdown. During normal system operation, these particulates (such as the used solid
catalyst) are carried outside of the device by the effluent. During shutdown, however, some
of these particulates are deposited in narrow areas inside the device. To circumvent this
problem in future experiments, it is proposed to add a compressed-air line to the setup, as
shown in Figure 6.1. The plug valve connected to this line would normally be turned off
during the experiments. At shutdown, this plug valve would be turned on while the other
plug valves connecting the peroxide and catalyst streams would be turned off. As a result,
the device could be purged, ensuring no deposition of particulates, and hence no clogging.
135
.
Chapter 6. Conclusions
PRV pressure relief.valve
Compressed
air line
PV Plug (on/off) valve
CV. Check valve
PRV
High-test
Peroxide
PV
rac;Kage
PRV
Product
waste
Ferrous
collector
chloride
tetrahydrateme
i illfil,
PV
Cat aly st
waste
collector
Figure 6.1. Schematic of proposed modified setup with compressed air line for purging the
device after experiments
A second problem that was encountered during experiments was the cracking of the
bottom half of the package along its side, as shown in Figure 6.2. The cracking was
observed to start in areas that were heavily stained by the catalyst. This cracking is
attributed to possible reaction of the package material with the used catalyst, which wetted
the bottom half of the package during startup. The brittle nature of the package's ceramic
material and the clamping mechanism are speculated to further aggravate this problem.
For future experiments, it is recommended, at least for the bottom half of the package, to
136
. ......
.-. .RO
............
. .....................
................
Chapter 6. Conclusions
use a less-brittle material that is inert to the catalyst. If finding a material that satisfies
these criteria while also having the required low thermal conductivity in Table 4.1 proves
to be challenging, a modified package design, such as the one shown in Figure 6.3 , can be
used. In this design, the extra cavities will be filled with air, which has a very low thermal
conductivity. This allows the use of a material with higher thermal conductivity than
Rescor 914.
-
5 mm
Figure 6.2. Photograph of the cracked bottom half of the package
Figure 6.3. Modified design for the bottom half of the package, which allows the use of a
material with higher thermal conductivity
137
Chapter 6. Conclusions
The third concern during experiments was safety. The incorporation of the safety
measures listed in Chapter 4 and the extreme care that was taken during experiments
prevented the occurrence of accidents. Nevertheless, it is more desirable to run future
experiments remotely, to minimize any potential risks. To achieve this, it is necessary to
automate the system. Automation can be realized by replacing the manual plug valves with
solenoid valves, and using an instrumentation software program, such as LABVIEW, to
control those valves and the syringe pumps during experiments. Extra piping lines (and
on/off valves) can be used to connect the syringes to separate reservoirs that contain the
peroxide and catalyst solutions. The instrumentation software can be used to run the
pumps in alternating withdrawal and infusion modes. During withdrawal, the syringes are
filled with the chemical solutions from the reservoirs. During infusion, the syringes supply
those solutions to the device. This prevents the need for manually filling the syringes after
each experiment.
6.3 Design Modifications for Future Applications
The device can be easily adapted for future inclusion in MEMS ejector pump systems.
The main design adjustment will be in the nozzle section. In the current device, the nozzle
emits gases into the atmosphere. In a pump application, the pump's upstream pressure,
which has to be matched to the steam generator's exit pressure, may be different. To allow
for a different exit pressure, the nozzle exit area can be increased or decreased. To first
order, the dimensions of the new required exit section can be found following the analysis
presented in Section 2.5.2 and using the new required pressure value, instead of 1 atm, for
Pe. The length of the supersonic section can be adjusted to avoid nozzle half-angles greater
138
Chapter 6. Conclusions
than about 300, which have been found to cause high viscous and non-axial flow losses in
computational studies by Louisos et al. [49-50]. For more compactness, the reactor length
may also be optimized to shorten the section following the location of complete peroxide
decomposition. When designing the ejector pump, attention must be paid to two important
points. First, some applications may require the pumped-out flow to be free of the catalyst
that was used for facilitating the peroxide decomposition. In these applications, care must
be taken to ensure that the pump design allows for the separation of the catalyst from the
flow, either under the action of gravity (as in the present steam generator) or using some
other mechanism. Secondly, the analysis presented in Section 2.9 can be used for the firstorder modeling of the pump's steady-state operation as a vacuum generator. The start-up
period, however, requires more involved modeling which is not covered here. During this
period, the pressure of the chamber to be evacuated decreases from an initial value to its
final desired value. The formation of shockwaves in the supersonic section of the steam
generator is expected during this stage in order to allow for higher-than-design evacuation
pressures in the suction chamber. Certain measures can be taken to reduce the energy
losses resulting from those shockwaves. The interested reader is referred to compressible
flow texts which address the starting of devices with supersonic flow and varying
downstream pressures, such as [51].
Another potential application that the device can be adapted for is thrust generation
in microrockets. The need for micropropulsion, or the production of very small thrust
levels, arises in two general areas of astronautics [45]. One area is the maintenance of very
precise orbital attitude and position in space missions. In this application, very low thrust
139
Chapter 6. Conclusions
levels (on the order of milli-Newtons) are required to offset perturbations that are caused,
for example, by solar radiation pressure or gravitational nonuniformities. Another area is
the miniaturization of space vehicles to facilitate the installation and use of distributed
spacecraft formations. The device presented in the current work can be adapted for use as
a microrocket driven by hydrogen peroxide as monopropellant. The thrust generated by
the steam-oxygen gas effluent can be utilized to propel the device. In micropropulsion, the
important performance metrics are the thrust force produced by the microrocket and its
specific impulse. The thrust force is proportional to the product of the propellant mass flow
rate and the flow velocity at the rocket exit. The specific impulse, which is a measure of the
thrust produced per unit weight of the propellant, is proportional to the flow velocity at the
exit. Therefore, to adapt the device for micropropulsion requirements, two changes can be
made to the design. The reactor geometry can be modified to enable the complete
decomposition of peroxide at a different mass flow rate. The supersonic section of the
nozzle can be made larger (wider at exit and longer), to allow effluent acceleration to higher
velocities as needed.
6.4 Concluding Remarks
We consider the research reported in this thesis to be another triumph in the areas of
micro-electro-mechanical systems and microfluidics. With detailed physical modeling,
sound, simultaneous design of both the device and the package, and thorough testing,
successful microsystems can be realized. Many of these systems are capable of providing
much-needed functionalities, and have scaling advantages in performance over their
macro-sized counterparts. The microsystem presented in this thesis is but one example.
140
Appendix A. Detailed process flow
Appendix A: Detailed Process Flow
Wafers 1-4
Lab
Process step
Machine
Alignment
marks
TRL
HMDS
TRL
Spin on resist
HMDSTRL
coater
Chemicals
OCG 825-
Color
Comments
green
thin resist recipe
green
1 ptm, positive resist
green
green
900C, 15 min
1 im, positive resist
green
green
90*C, 30 min
SteamGenAlignment
mask
20CS
(front)
oven
coater
TRL
TRL
Prebake
Spin on resist
(back)
Prebake
Expose front
TRL
Develop
photo-wet
TRL
Postbake
oven
green
900C, 30 min
ICL
Front oxide etch
AME5000
green
ICL
Front Si etch
AME5000
green
ICL
Piranha + rinse +
spin dry
ICLpremetal
green
Chamber A,0.5 jim
deep oxide etch
Chamber B, 0.25 pm
deep Si etch
piranha clean
TRL
HMDS
HMDS-
green
thin resist recipe
green
1 im, positive resist
green
green
90*C, 15 min
1 ptm, positive resist
TRL
TRL
OCG 82520CS
oven
EV1
OCG-934
sulfuric
acid,
peroxide
green
TRL
TRL
TRL
TRL
Spin on resist
(back)
prebake
Spin on resist
(front)
coater
oven
coater
OCG 82520CS
OCG 82520CS
TRL
Prebake
oven
green
900C, 30 min
TRL
Expose back
EVI
green
SteamGen._Alignment
mask
TRL
Develop
photo-wet
TRL
Postbake
oven
green
900C, 30 min
ICL
Front oxide etch
AME5000
green
ICL
Front Si etch
AME5000
green
ICL
Piranha + rinse +
spin dry
ICLpremetal
sulfuric
acid,
peroxide
green
Chamber A, 0.5 pim
deep oxide etch
Chamber B, 0.25 pm
deep Si etch
piranha clean
Wafer I
Lab
Process step
Machine
Chemicals
Color
Comments
Through
TRL
HMDS
HMDS-
green
thick resist recipe
green
12.5 pm, positive resist
OCG-934
green
TRL
holes
TRL
Spin on resist
coater
AZ4620
141
Appendix A. Detailed process flow
resist
(front)
TRL
Prebake
oven
TRL
Spin on resist
coater
AZ4620
green
900C, 30 min
green
10 ptm, positive resist
resist
(back)
TRL
Prebake
oven
green
90*C, 60 min
TRL
Expose
EV1
green
SteamGenHoles
TRL
Develop
photo-wet
TRL
Postbake
oven
TRL
Front oxide etch
(BOE)
Mount on quartz
acidhood/
ah2
coater
mask
AZ440
green
developer
TRL
BOE
green
90 0C, 30 min
green
0.5 ptm deep thermal
oxide etch
green
wafer
TRL
Si etch
green
STS2/3
560 im deep through
Si etch
TRL
Acetone dismount
photo-wet
acetone
green
acidhood/
ah2
sulfuric
acid,
green
piranha clean
Color
Comments
of wafer handle
TRL
Piranha + rinse +
spin dry
peroxide
Chemicals
Wafer 3
Lab
Process step
Machine
Oxidation
ICL
ICL
RCA clean
Oxide deposition
RCA
DCVD
green
green
TRL
Anneal
TubeB3
green
TRL
HMDS
HMDSTRL
green
4 jim, both sides
9500C, 1 hr
thick resist recipe
TRL
Spin on resist
(front)
prebake
Spin on resist
(back)
coater
green
12.5 pm, positive resist
green
green
900C, 30 min
10 gm, positive resist
Reactor +
Through
holes
TRL
TRL
oven
coater
AZ4620
resist
AZ4620
resist
TRL
Prebake
oven
green
900C, 60 min
TRL
Expose
EV1
green
SteamGenReactor
mask
TRL
Develop
photo-wet
TRL
Postbake
oven
green
90*C, 30 min
ICL
Front oxide etch
AME5000
green
4.5 pim deep oxide etch
TRL
Mount on quartz
coater
green
Si etch
STS2/3
green
Acetone dismount
photo-wet
acetone
green
acidhood/
ah2
sulfuric
acid,
green
AZ440
developer
green
wafer
TRL
TRL
940 pim deep through
Si etch
of wafer handle
TRL
Piranha + rinse +
spin dry
piranha clean
142
Appendix A. Detailed process flow
peroxide
Wafers 2
Color
Comments
RCA
DCVD
green
green
2 ptm, both sides of
Anneal
HMDS
TubeB3
HMDSTRL
green
green
9500C, 1 hr
thick resist recipe
Spin on resist
(front)
Prebake
Spin on resist
(back)
coater
green
12.5 tm, positive resist
green
green
900C,30 min
10 pm, positive resist
Lab
Process step
Machine
ICL
ICL
RCA clean
Oxide deposition
TRL
TRL
TRL
Chemicals
and 4 (Si)
Oxidation
each wafer
Mixers+rea
ctor+holes
in oxide
TRL
TRL
oven
coater
AZ4620
resist
AZ4620
resist
Prebake
oven
green
900C, 60 min
TRL
Expose front
EV1
green
SteamGenMixer
mask
TRL
Develop
photo-wet
TRL
ICL
ICL
Postbake
Front oxide etch
Piranha + rinse +
spin dry
oven
AME5000
ICLpremetal
TRL
HMDS
TRL
Spin on resist
(back)
HMDSTRL
coater
TRL
Prebake
oven
TRL
Spin on resist
(front)
coater
TRL
Prebake
TRL
TRL
AZ440
developer
green
green
green
green
900C,30 min
2.5 pim deep oxide etch
piranha clean
green
thick resist recipe
green
12.5 pm, positive resist
green
900C, 30 min
green
10 pm, positive resist
oven
green
900C, 60 min
Expose back
EV1
green
SteamGenMixer
mask
TRL
Develop
photo-wet
TRL
Postbake
oven
ICL
ICL
Back oxide etch
Piranha + rinse +
spin dry
AME5000
ICLpremetal
TRL
HMDS
TRL
Spin on resist
(front)
Prebake
Expose front
HMDSTRL
coater
TRL
TRL
oven
EV1
sulfuric
acid,
peroxide
AZ4620
resist
AZ4620
resist
AZ440
developer
sulfuric
acid,
peroxide
AZ4620
resist
green
green
90 0C, 30 min
green
green
2.5 pm deep oxide etch
piranha clean
green
thick resist recipe
green
12.5 jim, positive resist
green
green
900C, 30 min
SteamGenReactor
mask
143
Appendix A. Detailed process flow
photo-wet
TRL
TRL
Postbake
Back Si shallow
etch
Mount on quartz
wafer
Front Si deep etch
oven
STS2/3
green
green
coater
green
STS2/3
green
Oxygen plasma to
strip PR
Front Si shallow
etch
asher-TRL
green
STS2/3
green
photo-wet
acetone
green
TRL
Acetone dismount
of wafer handle
Piranha + rinse +
spin dry
acidhood/
ah2
sulfuric
acid,
peroxide
green
piranha clean
Lab
Process step
Machine
Chemicals
Color
Comments
TRL
TRL
Ash wafers
Piranha + rinse
green
green
1hour each wafer
piranha clean
TRL
Pure HF
green
5 min, for oxide
removal
TRL
EV620
green
TRL
TRL
RCA clean without
HF dip
Silicon direct
bonding
Pressing of stack
Anneal
TRL-asher
acidhood/ sulfuric
ah2
acid,
peroxide
acidhood/ hydrofloric
ah2
acid
RCA
EV501
TubeB3
green
green
9500C, 1 hr
Lab
Process step
Machine
Chemicals
Color
Comments
TRL
Piranha + rinse +
spin dry
acidhood
sulfuric
acid,
red
piranha clean
TRL
Anodic bonding of
Pyrex to Si +
EV501
red
1000 V for bonding
Diesaw
brown
TRL
TRL
Mixers+rea
ctor+holes
in Si
TRL
TRL
Wafers 1,
2. 3. 4 (all
TRL
Pyrex
wafer and
Si stack
green
Develop
TRL
Reactor +
holes in Si
AZ440
developer
TRL
900C, 30 min
100 ptm deep backside
etch
740 jim deep frontside
etch
100 ptm deep extra
frontside etch
green
peroxide
pressing
ICL
of stack
Dice wafers
144
Appendix B. Practices followed for successful wafer bonding
Appendix B: Practices Followed for Successful Wafer
Bonding
e
Order the silicon wafers with a 0.5 pm thick layer of thermally-grown oxide on each
side to protect the wafer surfaces during processing. The oxide layers are removed later
by using HF just prior to bonding.
" Use a new (unopened) wafer box and the RCA-designated boat (brown label) for
transporting wafers during the bonding process.
* The wafer box must be closed anytime it is being transported from one station to
another.
* It is important, prior to the RCA clean, to place the wafers in the boat according to the
order and orientation in which they will be bonded later. This is to avoid unnecessary
wafer handling afterwards to check for the correct wafer and side to be bonded.
* Following the HF cleaning, all subsequent steps must be carried out as quickly as
possible, and as soon as the previous step is over, with no waiting between steps. This is
to minimize the number of particulates from the cleanroom being deposited on the
wafers.
* Cover the counter space near the bonder with fabwipes.
* Clean any beakers/holders/wands that will be used during bonding with isopropanol.
" Wear sleeves and two sets of gloves, and have many extra (new) gloves handy near the
bonder station. Anytime you touch anything, the outer gloves must be replaced.
" Keep wafers in the spin-dryer while setting up the bonder and pressure chamber, and
do not bring the wafers to the bonder station until the setup task is fully completed.
This is to minimize the deposition of particulates on the wafers due to setup.
" Block access to the bonder station area (by other fab members) during bonding. Also
avoid talking with others and breathing directly on the wafers.
* Always run a pair of dummy wafers to catch any particulates present on the bonder
surfaces before bonding the actual wafers.
145
Appendix B. Practices followed for successful wafer bonding
*
To remove individual wafers with only one side to be bonded from the carrier, use the
vacuum wand on the side that will not be bonded.
146
References
References
1 Bogue, R.(2007). MEMS sensors: Past, present and future. Sensor Review, 27(1), 7-13.
2 Laser, D.J., &Santiago, J.G.(2004). A review of micropumps. Journalof Micromechanics
and Microengineering,14, R35.
3 Marcu B., Prueger G., Epstein A. H., Jacobson S. A. (2005). The commoditization of space
propulsion: Modular propulsion based on MEMS technology. 41st AIAA/ASME/SAE/
ASEE Joint Propulsion Conference & Exhibit.
4 Yao, S. C., Tang, X., Hsieh, C.C., Alyousef, Y., Vladimer, M., Fedder, G. K., et al. (2006).
Micro-electro-mechanical systems (MEMS)-based micro-scale direct methanol fuel cell
development. Energy, 31(5), 636-649.
5 Unger M. A., Chou, H. P., Thorsen, T., Scherer, A., & Quake, S. R. (2000). Monolithic
microfabricated valves and pumps by multilayer soft lithography. Science, 288(5463),
113.
6 Urbanski J.P., Thorsen, T., Levitan, J.A., & Bazant, M.Z. (2006). Fast AC electro-osmotic
micropumps with nonplanar electrodes. Applied Physics Letters, 89, 143508.
7 Edwards IV,J.M., Hamblin, M. N., Fuentes, H.V., Peeni, B.A., Lee, M.L., Woolley, A. T., et
al. (2007). Thin film electro-osmotic pumps for biomicrofluidic
applications.
Biomicrofluidics,1, 014101.
8 Van Lintel, H. T. G., Van de Pol, F. C. M., & Bouwstra, S. (1988). A piezoelectric
micropump based on micromachining of silicon. Sensors and Actuators, 15(2), 153-167.
9 Zimmermann, S., Frank, J.A., Liepmann, D., & Pisano, A. P. (2004). A planar micropump
utilizing thermopneumatic actuation and in-plane flap valves. 17th IEEE International
Conference on Micro Electro MechanicalSystems (MEMS 04), 462-465.
10 Zengerle, R., Ulrich, J., Kluge, S., Richter, M., & Richter, A. (1995). A bidirectional silicon
micropump. Sensors and ActuatorsA: Physical,50(1-2), 81-86.
11 Berg, J. M., Anderson, R., Anaya, M., Lahlouh, B., Holtz, M., & Dallas, T. (2003). A twostage discrete peristaltic micropump. Sensors and ActuatorsA: Physical,104(1), 6-10.
147
References
12 Dario, P., Croce, N., Carrozza, M. C., & Varallo, G. (1996). A fluid handling system for a
chemical microanalyzer. Journalof Micromechanicsand Microengineering,6, 95-98.
13 Benard W., Kahn, H., Heuer, A., & Huff, M. (1998). Thin-film shape-memory alloy
actuated micropumps. Journalof MicroelectromechanicalSystems, 7(2), 245-251.
14 Schomburg, W. K., Vollmer, J., Bustgens, B., Fahrenberg, J., Hein, H., & Menz, W. (1994).
Microfluidic components
in LIGA technique. Journal of Micromechanics and
Microengineering,4, 186-19 1.
15 Kamper, K. P., Dopper, J., Ehrfeld, W., & Oberbeck, S. (1998). A self-filling low-cost
membrane micropump. 11th IEEE InternationalConference on Micro Electro Mechanical
Systems (MEMS 98), 43 2-437.
16 B6hm, S., Olthuis, W., & Bergveld, P. (1999). A plastic micropump constructed with
conventional techniques and materials. Sensors &Actuators:A. Physical, 77(3), 223-228.
17 Cui, Z., & Takoudis, C. G. (2001). Design and simulation of a vacuum micropump.
Proceedingsof SPIE,4560, 263.
18 Vargo, S. E., Muntz, E. P., Shiflett, G. R., & Tang, W. C.(1999). Knudsen compressor as a
micro- and macroscale vacuum pump without moving parts or fluids.Journalof Vacuum
Science & Technology A: Vacuum, Surfaces,and Films, 17, 2308.
19 Vargo, S. E., &Muntz, E.P. (2001). Initial results from the first MEMS fabricated thermal
transpiration-driven vacuum pump. AIP Conference Proceedings,585, 502.
20 Young, R.M.(1999). Analysis of a micromachine-based vacuum pump on a chip actuated
by the thermal transpiration effect. Journal of Vacuum Science & Technology B:
Microelectronicsand NanometerStructures, 17, 280.
21 Hobson, J. P., & Salzman, D. B. (2000). Review of pumping by thermal molecular
pressure. Journal of Vacuum Science & Technology A: Vacuum, Surfaces, and Films, 18,
1758.
22 McNamara, S., & Gianchandani, Y. (2005). On-chip vacuum generated by a
micromachined Knudsen pump. Journalof MicroelectromechanicalSystems, 14(4), 741746.
148
References
23 Graham Corporation website.
Available: http://www.graham-mfg.com/index.asp?pageld=46
24 Croll Reynolds Company website.
Available: http://www.croll.com/ website/pr/vetheory.asp
25 Singhal, G.(2010). Pressure recovery studies on a supersonic COIL with central ejector
configuration. Optics &Laser Technology, 42(7), 1145-1153.
26 Doms, M., &Muller, J.(2007). Design, fabrication, and characterization of a micro vaporjet vacuum pump.Journalof FluidsEngineering,129, 1339.
27 Chuech, S. G., Chen, C.C., Lu, J. C., & Yan, M. M. (2007). Design and implementation of
ejector driven micropump. Energy Conversion and Management,48(10), 2657-2662.
28 Fan, Y., Suzuki, Y., & Kasagi, N. (2005). Development of large-entrainment-ratio micro
ejector for catalytic combustor. PowerMEMS 05, 28-30.
29 Chen, X., Zhang, T., Xia, L., Li, T., Zheng, M., Wu, Z., et al. (2002). Catalytic decomposition
of hydrazine over supported molybdenum nitride catalysts in a monopropellant
thruster. Catalysis Letters, 79(1), 21-25.
30 Timbretl, J. A., Scales, M. D. C., & Streeter, A. J. (1982). Studies on hydrazine
hepatotoxicity: 2. Biochemical findings. Journalof Toxicology and Environmental Health,
PartA, 10(6),955-968.
31 Ventura, M., & Mullens, P. (1999). The use of hydrogen peroxide for propulsion and power.
35th A IAA/ASME/SA E/ASEE Joint PropulsionConference & Exhibit.
32 Zimmerman, P. W. (1955). Chemicals involved in air pollution and their effect upon
vegetation. Contrib.Boyce Thompson Inst.
33 Weatherby, J.H. (1952). Chronic toxicity of ammonia fumes by inhalation. Proceedingsof
the Society for ExperimentalBiology and Medicine, 81(1), 300-301.
34 Scharlemann, C., Schiebl, M., Marhold, K., Tajmar, M., Miotti, P., Kappenstein, C., et al.
(2006). Development and test of a miniature hydrogen peroxide monopropellant
thruster. 42nd AIAA/ASME/SAE/ASEE Joint Propulsion Conference & Exhibit.
35 Yaws, C.L.(1999). Chemical propertieshandbook. New York, NY: McGraw-Hill.
149
References
36 Whitehead, J. C. (1998). Hydrogen Peroxide Propulsion for Smaller Satellites. 12th
Annual AIAA/USU Conference on Small Satellites.
37 Wilhite, B.A., Livermore, C., Gong, Y., Epstein, A. H., & Jensen, K. F. (2004). Design of a
MEMS-based microchemical oxygen-iodine laser (ptCOIL) system. IEEE Journal of
Quantum Electronics,40(8), 1041-1055.
38 Hill, T. F. (2008). Microchemical systems for singlet oxygen generation. Massachusetts
Institute of Technology, Cambridge, MA.
39 Dean, J., Aneshansley, D. J., Edgerton, H. E., & Eisner, T. (1990). Defensive spray of the
bombardier beetle: A biological pulse jet. Science, 248, 1219-1221.
40 Kuan, C. K., Chen, G. B., & Chao, Y. C. (2005). Development of a high-test hydrogen
peroxide (HTP) micro-thruster. 20th International Colloquium on the Dynamics of
Explosions and Reactive Systems.
41 Platt, D. (2002). A monopropellant milli-Newton thruster system for attitude control of
nanosatellites. 16th Annual USU Conference on Small Satellites.
42 Lee, J., An, S., & Kwon, S. (2008). Development of a liquid propellant rocket utilizing
hydrogen peroxide as a monopropellant. 44th AIAA/ASME/SAE/ASEE Joint Propulsion
Conference &Exhibit.
43 Musker, A.J., Roberts, G.T., Horabin, R., Kawar, A., Payne, M., Pollard, M., et al. (2006). An
assessment of homogeneous catalysts for the rapid decomposition of hydrogen
peroxide. 42nd AIAA/ASME/SAE/ASEE JointPropulsionConference & Exhibit.
44 Peroxide Propulsion website.
Available: http://www.peroxidepropulsion.com/article/14
45 Hitt, D.L., Zakrzwski, C.M., &Thomas, M.(2001). MEMS-based satellite micropropulsion
via catalyzed hydrogen peroxide decomposition. Smart Materials and Structures, 10(6),
1163-1175.
46 Xupeng, C., Yong, L., Zhaoying, Z., & Ruili, F. (2003). A homogeneously catalyzed microchemical thruster. Sensors &Actuators: A. Physical,108(1-3), 149-154.
47 Nguyen, N. T., &Wu, Z. (2005). Micromixers-a review. Journalof Micromechanics and
Microengineering,15(2), R1-R16.
150
References
48 Osaki, Y., & Takahashi, K. (2004). Microfluidics of liquid propellant microthruster for
pico-satellites. IEEJ Transactionson Sensors and Micromachines,123, 436-441.
49 Louisos, W. F., & Hitt, D.L. (2005). Optimal expansion angle for viscous supersonic flow
in 2-D micro-nozzles. 35th AIAA FluidDynamics Conference and Exhibit, 1-12.
50 Louisos, W. F., & Hitt, D. L. (2008). Viscous effects on performance of two-dimensional
supersonic linear micronozzles. Journal of Spacecraftand Rockets, 45(4), 706-715.
51 Greitzer, E. M., Tan, C.S., & Graf, M. B. (2004). Internalflow: Concepts and applications.
New York, NY: Cambridge University Press.
52 Musker, A. J., & Roberts, G.T. (2005). The effect of stabilizer content on the catalytic
decomposition of hydrogen peroxide. 41st AIAA/ASME/SAE/ASEE
Joint Propulsion
Conference & Exhibit.
53 Marco Rubber website.
Available: http://www.marcorubber.com/
54 Branebjerg, J., Gravesen, P., Krog, J.P., &Nielsen, C.R. (1996). Fast mixing by lamination.
9th IEEE InternationalConference on Micro Electro Mechanical Systems (MEMS 96), 441446.
55 Liu, R. H., Stremler, M. A., Sharp, K. V., Olsen, M. G., Santiago,
J. G., Adrian, R. J., et al.
(2000). Passive mixing in a three-dimensional serpentine microchannel. Journal of
MicroelectromechanicalSystems, 9(2), 190-197.
56 Wong, S. H., Ward, M. C.L., & Wharton, C.W. (2004). Micro T-mixer as a rapid mixing
micromixer. Sensors &Actuators: B. Chemical,100(3), 359-379.
57 Stroock, A. D., Dertinger, S. K.W., Ajdari, A., Mezic, I., Stone, H. A., & Whitesides, G. M.
(2002). Chaotic mixer for microchannels. Science, 295(5555), 647.
58 Yang, R., Williams, J. D., & Wang, W. (2004). A rapid micro-mixer/reactor based on
arrays
of
spatially
impinging
micro-jets.
Journal of
Micromechanics and
Microengineering,14, 1345-1351.
59 Mengeaud, V., Josserand, J., & Girault, H. H. (2002). Mixing processes in a zigzag
microchannel: Finite element simulations and optical study. Analytical Chemistry,
74(16), 4279-4286.
151
References
60 Wong, S. H., Bryant, P., Ward, M., & Wharton, C. (2003). Investigation of mixing in a
cross-shaped micromixer with static mixing elements for reaction kinetics studies.
Sensors &Actuators: B. Chemical, 95(1-3), 414-424.
61 Engler, M., Kockmann, N., Kiefer, T., & Woias, P. (2004). Numerical and experimental
investigations on liquid mixing in static micromixers. Chemical EngineeringJournal,
101(1-3), 315-322.
62 Kockmann, N., & Woias, P. (2003). Flow regimes and mass transfer characteristics in
static micromixers. ProceedingsofSPIE, 4982, 319.
63 Mays, L.W. (1999). Hydraulic design handbook. New York, NY: McGraw-Hill Professional.
64 Collins, K. C. (2008). Computationalfluid dynamic (CFD) optimization of microfluidic
mixing in a MEMS steam generator.Massachusetts Institute of Technology, Cambridge,
MA.
65 Mills, A.F. (1999). Heat transfer.Upper Saddle River, NJ: Prentice Hall Inc.
66 Zhou, X., & Hitt, D. L. (2003). One-dimensional
modeling of catalyzed H20 2
decomposition in microchannel flows. 33rd AIAA FluidDynamics Conference and Exhibit.
67 Orr, R. J., & Williams, H. L. (1953). Kinetics of the reactions between iron (II) and
hydroperoxides based upon cumene and cyclohexane. The Journalof Physical Chemistry,
57(9), 925-931.
68 Kundu, P. K., & Cohen, I. M. (2004). Fluid mechanics. San Diego, CA: Elsevier Academic
Press.
69 FMC Corporation website.
Available: http://www.fmcchemicals.com/Products/HydrogenPeroxide.aspx
70 Bernier, M. (2004). An experimental investigation of heat transferto hydrogen peroxide in
microtubes.Massachusetts Institute of Technology, Cambridge, MA.
71 Quinn, W. R. (1994). Development of a large-aspect-ratio rectangular turbulent free jet.
AIAA Journal,32(3), 547-554.
152