Design of a Fast Cycle Time Hot Micro-Embossing Machine
By
Matthew E. Dirckx
B.S., Mechanical Engineering
University of Oklahoma, 2003
Submitted to the Department of Mechanical Engineering
in Partial Fulfillment of the Requirements for the Degree of
Master of Science in Mechanical Engineering
at the
Massachusetts Institute of Technology
June 2005
MASSACHUSETTS INS
OF TECHNOLOGy
JUN 16 2005
D Massachusetts Institute of Technology
All Rights Reserved
Signature of Author .....................
C ertified by
...........
LIBRARIES
................................
Department of Mechanical Engineering
May 6, 2005
.
. . . . . . . . . . . . . . . . .
.
David E. Hardt
Professor of Mechanical Engineering
Thesis Supervisor
Acceptedby.......................
..............................
Lallit Anand
Chairman, Department Committee on Graduate Students
BARKER
1
E
Design of a Fast Cycle Time Hot Micro-Embossing Machine
by
Matthew E. Dirckx
Submitted to the Department of Mechanical Engineering on May 6, 2005
in Partial Fulfillment of the Requirements for the Degree of
Master of Science in Mechanical Engineering
ABSTRACT
In the coming years, there will be a huge market for mass-produced polymer microdevices. These devices include microfluidic "labs on a chip," micro-optical chips, and
many others. Several techniques exist for producing micron-scale features in polymer
materials. One of the most promising of these techniques is Hot Micro-Embossing
(HME). In this process, a thermoplastic polymer workpiece is heated above its glass
transition temperature and a micro-patterned die is forced into it. The polymer conforms
to the workpiece and the features are replicated. Much of the research to date concerning
HME has not addressed fundamental issues that will be central to successful mass
production using this process. There is a compelling need to study HME from the
perspective of manufacturing process control. In order to conduct such a program, a
HME machine is needed that allows the operator to precisely control all the potentially
significant process parameters. No existing machine fully meets this requirement. This
thesis concerns the conceptual and detailed design of a HME system, including the platen
assembly and the temperature control system. A parametric model and finite element
analysis were used to guide the design of the platen assembly and to assess its thermal
and structural performance. A dynamic thermal model of the temperature control system
was developed. This model was used to guide the selection of components and to predict
the performance of the system as a whole. The new design will have a short cycle time,
will permit the use of full wafer-size embossing tools, and will be able to follow a userprogrammed trajectory in displacement, force, and temperature.
Thesis supervisor: David E. Hardt
Title: Professor of Mechanical Engineering
2
Acknowledgements
First, I wish to thank my parents for their love and guidance. With their unending
support and encouragement, and by their example, I have gone far.
I would like to thank my advisor, Prof. Dave Hardt, for the opportunity to join his
lab, and for his mentorship during my work.
I would also like to thank my colleagues. Grant Shoji and Kunal Thaker directly
contributed to this work by their research into heat transfer fluids, control valves, vacuum
pumps, and hot oil pumps, by their work to select and procure these and many other
components of the fluid system, and by their analysis of the system pressure drop.
Without their assistance in these and many other areas, the design of the new machine
would have been a much greater burden. Adam Rzepniewski and Wang Qi have also
contributed their advice, and Catharine Nichols has been swamped with paperwork for
this project. All of my colleagues have made the lab a friendly and fun place to work.
I would also like to thank the staff of the LMP machine shop, especially Gerry
Wentworth, for all of their instruction and help with machining the components for the
new machine.
Tricia has given me her love and support, and has endured the crowds and climate
of Boston, and for this, I will always be grateful.
Finally, I would like to thank the Singapore-MIT alliance for funding this work
and making my studies at MIT possible.
3
Table of contents
A cknow ledgem ents.......................................................................................................
Table of contents.................................................................................................................
List of figures......................................................................................................................
N omenclature ......................................................................................................................
Introduction...............................................................................................................
1
G oing "M icro" ................................................................................................
1.1
M anufacturing at the m icron scale................................................................
1.2
Overview of thesis ............................................................................................
1.3
Background ...............................................................................................................
2
Techniques for micron-scale polym er replication .........................................
2.1
Soft lithography ....................................................................................
2.1.1
M icro-injection molding .......................................................................
2.1.2
U ltraviolet em bossing ...........................................................................
2.1.3
H ot m icro-em bossing............................................................................
2.1.4
Prior work in HME manufacturing process control.......................................
2.2
The m anufacturing process control paradigm ................................................
2.3
Existing H M E m achines ................................................................................
2.4
The generation 1 H ME machine ...........................................................
2.4.1
Com m ercially available H M E m achines ...............................................
2.4.2
The need for a new HM E m achine ...................................................................
2.5
G oals for the new m achine .......................................................................................
3
Introduction...................................................................................................
3.1
Probing spatial variation ................................................................................
3.2
Workpiece m aterial and therm al requirem ents .............................................
3.3
Tim e-dom ain process param eters ..................................................................
3.4
A utomation ....................................................................................................
3.5
W orkpiece and tool fixturing .........................................................................
3.6
Project scope ..................................................................................................
3.7
Summary of goals for the generation 2 HME machine .................................
3.8
Concept developm ent and evaluation ....................................................................
4
Introduction...................................................................................................
4.1
Tem perature control.......................................................................................
4.2
Therm oelectric (Peltier).........................................................................
4.2.1
ixed electric & fluid...............................................................................
M
4.2.2
A ll fluid...................................................................................................
4.2.3
Concepts for therm al fluid supply system ....................................................
4.3
Bulk heating & cooling.............................................................................
4.3.1
Separated stream s...................................................................................
4.3.2
Selection of w orking fluid .............................................................................
4.4
W orkpiece and tool fixturing .........................................................................
4.5
Sum mary of conceptual design.....................................................................
4.6
5
D esign of the platen assem bly ..............................................................................
Introduction.....................................................................................................
5.1
4
3
4
7
9
11
11
12
13
14
14
14
15
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19
21
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39
40
41
43
43
43
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45
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47
48
48
51
52
53
55
55
D esign of the platens.....................................................................................
5.2
Param etric m odel of platen perform ance..............................................
5.2.1
Manufacturability & selection of tube diameter ..................
5.2.2
Specifying a flow rate ...............................................................................
5.2.3
Final platen design ................................................................................
5.2.4
Fixturing and mounting.................................................................................
5.3
M ounting the platens..............................................................................
5.3.1
W orkpiece clamp ..................................................................................
5.3.2
Tool chuck (vacuum )..............................................................................
5.3.3
Spacer plate............................................................................................
5.3.4
Structural and Thermal finite element model of platen assembly ................
5.4
Structural.................................................................................................
5.4.1
Therm al..................................................................................................
5.4.2
M anifolds .......................................................................................................
5.5
Insulation.......................................................................................................
5.6
Summ ary of platen assembly design..............................................................
5.7
D esign of the temperature control system ..............................................................
6
Introduction.....................................................................................................
6.1
Selection of oil/w ater heat exchanger .............................................................
6.2
D ynam ic therm al m odel..................................................................................
6.3
Platen model............................................................................................
6.3.1
H eater m odel...........................................................................................
6.3.2
Oil/w ater heat exchanger model .............................................................
6.3.3
Control valve model................................................................................
6.3.4
Selection of electric circulation heater............................................................
6.4
Predicted dynam ic therm al perform ance ........................................................
6.5
M odeling fluid losses......................................................................................
6.6
Selection of control valves......................................................................
6.6.1
Selection of pump & m otor.....................................................................
6.6.2
Sizing of the Expansion Tank .........................................................................
6.7
Safety equipm ent ............................................................................................
6.8
Summary of Temperature control system design ....................
6.9
7
Conclusions and future w ork ..................................................................................
Sum m ary........................................................................................................
7.1
The final design.......................................................................................
7.1.1
Predicted system perform ance ................................................................
7.1.2
7.2
Conclusion ......................................................................................................
Future w ork.....................................................................................................
7.3
Appendix.........................................................................................................................
M aterial properties ..........................................................................................
A
Properties of Paratherm MR .......................................................................
A .1
Properties of PMM A ...................................................................................
A .2
Properties of Copper ...................................................................................
A .3
Properties of Silicon....................................................................................
A .4
Properties of Therm agon T-Pli 220 ............................................................
A .5
Properties of Rescor 914 Glass Ceram ic ....................................................
A .6
5
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57
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65
71
71
72
74
75
77
78
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81
87
94
95
100
100
101
104
105
107
108
108
109
112
119
120
122
124
126
127
128
128
128
129
129
130
132
133
133
135
138
139
141
143
Com ponent drawings ......................................................................................
144
B.1
B.2
Platen...........................................................................................................
Spacer..........................................................................................................
144
145
B.3
Clam p..........................................................................................................
146
B
V acuum chuck ............................................................................................
M anifold......................................................................................................
Bottom carrier plate ....................................................................................
Top Carrier..................................................................................................
Screw block.................................................................................................
T-slotted table .............................................................................................
Platen assembly...........................................................................................
Oil-W ater heat exchanger ...........................................................................
H eater..........................................................................................................
M atLab code ...................................................................................................
C
Properties of Paratherm M R .......................................................................
C.1
C.2
Param etric m odel of internal convection for platens ..................................
B.4
B.5
B.6
B.7
B.8
B.9
B.10
B. 11
B.12
147
148
149
150
150
151
152
153
154
156
156
156
Dynam ic therm al m odel..............................................................................
157
C.3.1
M ain program ......................................................................................
157
C.3.2
C.3.3
Cold heat exchanger module...............................................................
Electric heater m odule ........................................................................
159
159
C.3.4
Platens .................................................................................................
160
C.3.5
Calculate branch flows........................................................................
161
C.3
References.......................................................................................................................
6
163
List of figures
Figure 2-1 Schematic for Soft Lithography..................................................................
Figure 2-2 Schematic for micro-injection molding ......................................................
Figure 2-3 Schematic for UV embossing.....................................................................
Figure 2-4 Schematic for hot micro-embossing...............................................................
Figure 2-5 Temperature and force trajectory in HME ..................................................
Figure 2-6 Generic manufacturing process model.......................................................
Figure 2-7 Generation 1 machine overview................................................................
Figure 2-8 Generation 1 machine platens ....................................................................
Figure 2-9 Table of commercial hot embossing machines ...........................................
Figure 2-10 EV Group 520HE....................................................................................
Figure 2-11 Obducat NIL-4 .........................................................................................
Figure 2-12 Jenoptik HEXOI......................................................................................
Figure 2-13 Suss SB 6e..................................................................................................
Figure 3-1 Thermal model of workpiece ......................................................................
Figure 3-2 Simplified thermal model of workpiece with boundary conditions...........
Figure 4-1 Bulk heating & cooling of fluid ..................................................................
Figure 4-2 Separated hot & cold streams....................................................................
Figure 4-3 Photo of an electric circulation heater []...................................................
Figure 4-4 Diagram of plate and frame heat exchanger [].................
Figure 5-1 Basic schematic of platen assembly...........................................................
Figure 5-2 Minimal platen design................................................................................
Figure 5-3 Convection coefficient results from parametric model...............................
Figure 5-4 Dependence of platen mass on tube diameter .............................................
Figure 5-5 Pressure drop from parametric model.........................................................
Figure 5-6 Volume flow rate from parametric model..................................................
Figure 5-7 Convection coefficient vs. flow velocity ....................................................
Figure 5-8 Pressure drop vs. flow velocity ..................................................................
Figure 5-9 Biot number vs. flow velocity....................................................................
Figure 5-10 Time constant vs. flow velocity ...............................................................
Figure 5-11 Final platen design ....................................................................................
Figure 5-12 T-slotted table............................................................................................
Figure 5-13 B ottom carrier plate......................................................................................
Figure 5-14 Top carrier plate mounted to anvil...........................................................
Figure 5-15 Workpiece clam plate...................................................................................
Figure 5-16 Vacuum chuck.........................................................................................
Figure 5-17 Detail of vacuum port...................................................................................
Figure 5-18 The spacer plate.........................................................................................
Figure 5-19 Detail of structural FEA model of platen..................................................
Figure 5-20 von Mises stress in platen (Pa)..................................................................
Figure 5-21 Vertical deflection at top of platen...........................................................
Figure 5-22 Platen stack model.....................................................................................
Figure 5-23 Temperature in center of PMMA .............................................................
Figure 5-24 Temperature Distribution at bottom of PMMA over time .......................
7
15
16
17
18
19
21
24
25
27
27
27
28
28
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34
48
49
50
51
56
57
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61
62
63
66
67
68
70
71
73
73
74
75
76
77
78
79
80
81
82
82
84
Figure 5-25 Edge effect over time ................................................................................
Figure 5-26 Fluid temperature change along a tube ....................................................
Figure 5-27 Flotran CFD model of manifold design ....................................................
Figure 5-28 Mean flow velocity in each tube in the CFD model .................................
Figure 5-29 Convection coefficient in each tube.........................................................
Figure 5-30 Manifold design ......................................................................................
Figure 5-31 Stress distribution in the manifold ...........................................................
Figure 5-32 Thermal stress model ................................................................................
Figure 5-33 Exploded view of platen stack ..................................................................
Figure 5-34 The full platen assembly ...........................................................................
Figure 5-35 Three-dimensional view of the full platen assembly ................................
Figure 5-36 Table of thermal masses (*=estimated property).......................................
Figure 6-1 System A rchitecture .....................................................................................
Figure 6-2 Mixing hot and cold fluid to produce desired temperature ..........................
Figure 6-3 Photo of oil cooler........................................................................................
Figure 6-4 Oil cooler performance ................................................................................
Figure 6-5 Information flow diagram for the dynamic thermal model............
Figure 6-6 Photo of circulation heater ...........................................................................
Figure 6-7 Output of dynamic thermal model with the 30kW circulation heater.....
Figure 6-8 Performance with 25kW heater....................................................................
Figure 6-9 Performance with a 35kW heater.................................................................
Figure 6-10 Temperature dynamics of other system components ...............
Figure 6-11 Hot and cold branch flow rates ..................................................................
Figure 6-12 Temperature dynamics for 90-120'C step .................................................
Figure 6-13 Temperature dynamics for 30-135-60'C steps .........................................
Figure 6-14 R am p response ...........................................................................................
Figure 6-15 R am p tracking error ...................................................................................
Figure 6-16 Table of component pressure drops (kPa)..................................................
Figure 6-17 Control valve and positioner .......................................................................
Figure 6-18 Pump, gearbox, and motor .........................................................................
Figure 6-19 Table of component volumes.....................................................................
Figure 6-20 Diagram of expansion tank with cold trap .................................................
8
85
87
88
89
90
91
92
93
96
97
98
99
101
102
103
103
105
109
110
111
112
114
115
116
117
118
119
120
122
124
125
126
Nomenclature
Symbol
Ra
Re
SG
SV
T
t
T,,
Tc
Description
Sensitivity of the output of a process model to disturbances
Sensitivity of the output of a process to changes in the input
Area
Biot number
The nth coefficient for the Fourier sine series
An arbitrary constant
Specific heat capacity
Valve flow coefficient
Diameter
Effective diameter for fluid flow
Internal diameter of the shell of a shell and tube heat exchanger
Elastic modulus
Fluid friction factor
Gravitational acceleration
Grashof number
Convection coefficient
Fluid head
Thermal conductivity
Thermal expansion coefficient
Length
A positive integer
Number of baffles in a shell and tube heat exchanger
Nusselt number
Pressure drop across valve
Total system pressure drop not including valves
Prandtl number
Heat transfer rate
Volume flow rate
Rate of Joule heating in heating elements (Heater power)
Rayleigh number
Reynolds number
Specific Gravity
System volume not including expansion tank
Temperature
Time
Ambient temperature
Temp of cold fluid stream
Td
Desired temperature
Tg
Glass transition temperature
Temp of hot fluid stream
Temperature of fluid
Thermal mass
aY/aa
aY/3u
A
Bi
B11
C
CP
Cv
D
De
Ds
E
f
g
Gr
h
H
k
K
L
n
Nb
Nu
P1
P2
Pr
q
Q
qE
TI,
Tm
TM
9
Tp
Ts
Tt
V
VVA
x
XT
z
a
P
Aa
Ap
Au
AY
C
o
0,
p
v
p
T
0D
Temperature of platen
Temperature of surface subject to convection
Temperature of heating elements
Flow velocity
Volume
Valve Authority
Distance
Expansion tank volume
Fluid height difference
Thermal diffusivity
Gas expansion coefficient
Disturbances to a process model
Pressure drop across a component
Changes to the input of a process model
Changes in the output of a process model
Elastic strain
Temperature difference
Initial temperature difference
Dynamic viscosity
Kinematic viscosity
Density of a material
Exponential time constant
Dimensionless temperature=/0i
10
CHAPTER
1
Introduction
1.1 Going "Micro"
Progressive miniaturization has been one of the defining themes of technological
advance throughout the past century. Miniaturization and integration of electronic
components on microchips have enabled exponential increase in computers' capability
and utility. Toward the end of the
2 0 th
Century this paradigm was applied to biochemical
processes as heralded in 1990 by a key paper by Manz and Widmer [1]. They envisioned
miniaturizing and integrating all of the components of complex biochemical systems onto
a single microfluidic chip, directly analogous to microelectronic computer chips. Early
developers of microfluidic technology drew on the well-established methods of
micromachining of silicon and glass to produce micron-scale channels, chambers, and
other fluidic components [2,3]. Although some companies successfully commercialized
microfluidic devices using legacy microelectronics manufacturing methods in glass and
silicon [4], there is a growing consensus that the future of microfluidics lies in cheap,
disposable products manufactured from polymer materials [5,6,7].
Several processes exist that are capable of producing micron-scale features in
polymers, including material removal techniques such as laser ablation and X-ray
lithography [5], as well as forming and replication methods such as micro-injection
molding, soft lithography, ultraviolet embossing, and hot micro-embossing [8]. Many of
these processes are equally capable of producing micron-scale optical features such as
lenses [9]. Most of these processes have only recently been applied at this scale or are
11
entirely new. Consequently, the base of manufacturing know-how that permitted quick
commercialization of silicon and glass microfluidics does not exist.
1.2 Manufacturing at the micron scale
Credible estimates of the market for polymer micro-devices are on the order of
billions of dollars [6], so there exists a compelling need to develop a scientific
understanding of the technology for creating these devices from a manufacturing
perspective. Such a view embraces a broad array of concerns including process
capability, productivity, and quality; however, the most salient feature of manufacturing
as distinct from "one-off' prototyping is the concern for variation in the final product.
Manufacturing process control seeks to confront this variation by identifying its sources,
analyzing the interaction of the material and the machine, and applying methods such as
statistical process control and feedback control to reduce variation to acceptable levels.
Every manufacturing endeavor involves some degree of process control if it is to
be successful, and indeed the field of process control for established technologies such as
casting and machining of metals and macro-scale molding of polymers is well developed.
Manufacturing at the micron scale presents several unique challenges including material
purity, precise machine control, and high-resolution metrology. Many of these issues
have been addressed for micromachining of silicon and glass in the microelectronics
industry; however, such a foundation has yet to be established for the comparatively new
processes that will produce the polymer micro-devices of the future.
12
1.3 Overview of thesis
The Manufacturing Process Control Laboratory (MPCL) at the Massachusetts
Institute of Technology has a long history of applying the process control toolbox to
novel technologies such as discrete-die sheet metal forming and gas-metal-arc welding.
The current study is part of a larger research program whose objective is the development
of a foundation of knowledge for the application of manufacturing process control to hot
micro-embossing for the production of polymer micro-devices.
This thesis concerns the design of a machine for hot micro-embossing (HME) to
be used in the MPCL. Chapter 2 surveys prior work in the area of manufacturing process
control for hot micro-embossing as well as existing embossing machines, including one
that was previously developed in the MPCL. While these machines are quite capable, a
new machine is needed to advance the study of process control for HME, and chapter 3
discusses the requirements and goals for this new machine. Chapter 4 concerns the
development and evaluation of conceptual designs for the new machine, as well as the
selection of the final concept. Chapters 5 and 6 present the subsequent detailed design
and analysis of the platen assembly and the temperature control system respectively.
Chapter 7 summarizes the final design and predicted performance of the machine, and
discusses conclusions from this study and suggestions for future work.
13
CHAPTER
2
Background
2.1 Techniques for micron-scale polymer replication
Hot Micro-Embossing (HME) is one of a family of replication techniques for
producing micron-scale features in polymers. Related processes include soft lithography,
micro-injection molding, and ultraviolet embossing. These processes are discussed in
depth below.
2.1.1 Soft lithography
Soft lithography is a method for replicating micro-features by net-shape casting of
elastomeric polymers. The typically material of choice is polydimethylsiloxane (PDMS),
although other materials have been used [10]. PDMS has the advantages of being
optically transparent, thermally stable, and biocompatible. The basic process schematic
is presented in figure 2-1. In step 1, a mixture of PDMS resin and a curing agent is
poured over a micro-patterned tool. These tools are typically made using traditional
lithographic techniques, and may be re-used to produce several elastomer replicates In
step 2, the resin mixture conforms to the tool shape and cures over. Curing time depends
on the ratio of resin to curing agent and the ambient temperature, but is typically on the
order of hours. In step 3, the cross-linked elastomer is removed from the tool, usually by
carefully peeling the flexible structure off by hand. Soft lithography is uniquely capable
of producing micro-features in elastomers such as PDMS, but other processes must be
14
used if a rigid product is desired. The process cycle time is also inherently limited by the
curing time.
10
3
.r...
Figure 2-1 Schematic for Soft Lithography
2.1.2 Micro-injection molding
Micro-injection molding is similar its macro-scale namesake. A basic schematic
is shown in figure 2-2. In step 1, a molten polymer is forced into a mold cavity
containing a micro-structured insert. In step 2, the polymer cools rapidly. In step 3 the
mold is opened and the polymer is removed. Cycle times are typically very short, as the
small volumes of polymer lose heat to the large metal mold structure and cool rapidly.
Su et al found that the final quality of micro-injection molded parts is very sensitive to
process parameters, especially mold temperature [11]. The large amount of bulk material
15
flow as well as the large thermal cycle tends to produce residual stresses and shrinkage in
the final part.
34
Figure 2-2 Schematic for micro-injection molding
2.1.3 Ultraviolet embossing
In ultraviolet embossing, shown in figure 2-3, a UV-curable epoxy resin is applied
between a substrate and a micro-patterned tool. In step 2, the tool is brought into contact
with the resin and the resin conforms to its shape. The resin is exposed to a UV light
source and cured. In step 3, the polymer part is removed from the tool. Rossi and
Kallioniemi describe the application of this process to produce precise micro-lenses [9].
This process can only be used with UV-curing resins, and either the tool or the substrate
must be transparent to UV radiation.
16
2
Figure 2-3 Schematic for UV embossing
2.1.4 Hot micro-embossing
In HME, a thermoplastic polymer is formed by visco-plastic deformation. A
schematic of the process is shown in Figure 2-4. The polymer workpiece and the micropatterned tool are initially at ambient temperature. In step 1 the workpiece and tool are
heated above the glass transition temperature. In step 2, a forming force is applied and
held constant for a time to force the polymer to conform to the tool. In step 3, the
polymer and tool are cooled together, and then the polymer is removed from the tool.
Hot embossing has the advantage that the thermal cycle is smaller and bulk material flow
is reduced, minimizing residual stress and shrinkage. Another view of the HME process
is given in figure 2-5. The figure shows the temperature and force trajectory over time.
17
Some of the important process parameters are visible from this figure. The workpiece
temperature during embossing, the embossing force, and the duration of hold time are
presumed to affect the polymer's ability to conform to the tool. De-embossing
temperature may have an effect on any distortions introduced at that time. In addition to
dominating the total cycle time, the rate of heating and cooling may also have an effect
on final quality. Understanding the nature and significance of the relationships between
these process parameters and final quality is the first step in developing a foundation for
manufacturing process control for HME.
2
3
Figure 2-4 Schematic for hot micro-embossing
18
Force
Time
Temperatureold
-e-p-r-u-e-.-.-.Em bossing
Force
Embossing
Temperature
Tg
-- .-.-
-
-
---
..
De-Embossing
Temperature
Ambient
Temperature
0
...
De-Embossing
Force
Time
Figure 2-5 Temperature and force trajectory in HME
2.2 Prior work in HME manufacturing process control
Hot micro-embossing has been used by several researchers to produce microfeatures in thermoplastics, especially polymethylmethacrylate (PMMA). The
preponderance of publications on this subject concerns production of proof-of-concept
devices, while relatively few have addressed issues relevant to process control [12].
Roos et al used a commercially available hot embossing machine to conduct some
studies of the HME process. They used a EVG-520 wafer bonding machine modified for
hot embossing to imprint 100mm wafers coated with a thin layer of PMMA, and
qualitatively evaluated the effect of varying embossing temperature and force [13]. In a
later paper they evaluate the difference in quality when embossing is performed at
atmospheric pressure vs. under a vacuum, finding that vacuum improved uniformity over
the part [14]. Similarly, Bacon et al used the same type machine and compared the
results of 49 combinations of temperature and embossing force [15]. In none of these
papers was any effort made at rigorous design of experiments, nor was any statistical
19
analysis done or any attempt made to derive a mathematical process model from the data.
None of these studies considered any time-domain parameter such as hold time, strain
rate, heating rate, or cooling rate. Lin et al compared the quality of embossed parts made
in a laboratory process with those made in a commercial process and found that the
laboratory process replicated features better [16]. Significant to this discussion is the fact
that the laboratory process took about two hours, while the commercial process took only
a few minutes. Direct quantitative comparison is not possible, because the lab and
commercial processes used different workpiece materials and different tool materials, but
this study at least suggests that time-domain parameters are significant.
A significant addition to the literature on HME process control was made by
Ganesan in his SM thesis [17]. He designed and built a lab-scale hot embossing machine
and used this machine to investigate the natural variability of the HME process.
Processing parameters were held constant and several PMMA parts were formed using an
etched silicon die with channels and other features ranging 3 microns to 170 microns
wide and 1 micron deep. The resulting features in the PMMA were measured using an
optical profilometer. The sizes of various features were compared to the size of
corresponding features on the tool. The standard deviation of part dimensions as well as
the magnitude of the die-part difference were found to be on the order of the
measurement resolution of about 0.5 microns, and were found to scale strongly with
feature size. Statistical process control charts for the feature dimensions were also
analyzed, and some features were found not to be in a statistical state of control. These
deviations were largely confined to certain areas of the parts, and were attributed to a lack
of precise control over certain process parameters including cooling.
20
Later, Thaker, Shoji, et al. re-measured Ganesan's sample parts using an atomic
force microscope [18]. The higher resolution of this method allowed them to better
characterize the dimensional variation in the smaller-size features of the parts. They
found that for raised features about 1 micron tall and an average of 4.08 microns wide,
the standard deviation of width was 0.52 microns, or 12.7% of the dimension. For most
manufacturing processes, a 12% deviation in the final part would be unacceptable.
Suggested causes for this level of variation included variability in the workpiece material
and cycle to cycle variability of process parameters.
2.3 The manufacturing process control paradigm
There is a clear need to develop a basic, thorough understanding of hot microembossing from a manufacturing process control perspective. All manufacturing
processes involve the application of energy to transform a material. In the case of HME,
this consists of thermal energy to heat the polymer, mechanical work to force the polymer
to conform to the die, and thermal energy to cool the polymer. A generic process model
was introduced by Hardt and is presented in figure 2-6 [19]. The user provides the
process with certain inputs, and the machine applies energy to the material to produce the
desired outputs, subject to disturbances.
Disturbances
a
Energy
Machine -Material
Inputs
Outputs
(Part Geometry)
Figure 2-6 Generic manufacturing process model
21
Mathematically, the resulting variation may be approximated by equation 2-1.
For the single input-single output case, a change in the output is given by AY, which is a
function of the parameter disturbances Aa multiplied by the sensitivity of the output to
such disturbances, DY/Da, added with the change in the input Au multiplied by the
sensitivity of the output to such changes (the input-output gain) Y/u.
DY
AY=
DY
Aa+
Au
Equation 2-1
For a linear multi-input multi-output process, the disturbance term and the input
term would be vectors, and the sensitivity terms would be matrices. Off-diagonal
elements in these sensitivity matrices would represent coupling between different
parameters or disturbances. In the most generic case, this relationship could be nonlinear
as well.
Once the parameters of the generic process model are known, manufacturing
process control can be applied. One may attempt to reduce parameter disturbances
through clever machine design and control, or compensate for disturbances by changing
the inputs. With knowledge of the sensitivity functions (or matrices), one can select
''optimal" values for the parameters that produce minimum values for the sensitivity
function, thus minimizing the sensitivity of the output to disturbances.
2.4 Existing HME machines
All efforts at optimizing the hot micro-embossing process or at implementing
manufacturing process control depend on a thorough understanding of the process. In
other words, the terms of the process model equation must be known for the significant
process parameters and inputs. While some have investigated a limited subset of process
22
parameters such as embossing force and embossing temperature, no purposeful effort has
yet been made to examine the full range of putative process parameters for significance
or to quantify the relationships among parameters, inputs, and outputs. If an embossing
apparatus is to be adequate for such experiments, it must be capable of precise control of
the process parameters experiments throughout their practical ranges.
2.4.1 The generation 1 HME machine
Ganesan designed and built a capable HME machine for his SM thesis at the
MPCL [17]. This machine has proven useful both for his experiments and for others. An
Instron model 5869 electromechanical load frame provides force and position control for
the embossing platens. The platens themselves are blocks of copper heated electrically
with cartridge heaters and cooled by tap water. The heaters were controlled using
Chromalox 2110 controllers with temperature feedback. This apparatus is shown in
figure 2-7. A close-up view of the platens and workpiece fixturing is shown in figure
2-8. The heater wires and cooling tubes can also be seen in this figure.
23
Figure 2-7 Generation
24
1 machine overview
Figure 2-8 Generation 1 machine platens
The generation 1 machine is capable of embossing forces up to 50 kN and
temperatures up to about 300'C. Displacement of the crosshead can be controlled with a
resolution of 0.0625 [tm up to a speed of 250 mm/min. This capability allows the
machine to follow arbitrary position or force trajectories within its limits, so the full
range of process parameters in the position and force domains can be investigated. These
parameters include embossing force, embossing strain rate, hold force, maximum tool
displacement, and others.
Embossing temperatures up to 300'C can be set to ±1C, and arbitrary deembossing temperatures can be chosen, however the user does not have control of heating
rate. Cooling is controlled with manual water valves, and so is not very repeatable or
25
controllable. De-embossing temperature is not controlled with precision, since the flow
of water is shut off manually. Large copper blocks were needed for the platens in order
to ensure even heat distribution to the workpiece and tool, and their large thermal mass
limited the speed of temperature change. Typical experiments involve heating from
ambient to 130'C or higher, taking about 15min. Cooling back to ambient is more rapid,
taking about 5min. The largest workpiece that can fit in the fixturing area is about 45mm
by 40mm, although in practice this has been limited to about 25mm. The small size of
these test pieces served to eliminate special variation of process parameters across the
workpiece and tool to eliminate a confounding factor in experiments. The tool was
affixed using high-temperature epoxy to a post mounted to the upper platen. The
workpiece was clamped around its periphery by a copper plate with a hole through which
the tool post could pass.
A thorough study of the HME process will require control over the temperature
and temperature trajectory throughout the embossing process similar to the existing
degree of control over force and displacement. Indeed, Ganesan notes that better control
of process parameters is necessary to advance the level of understanding for HME [17].
Faster heating and cooling is also needed to investigate the lower end of these process
parameters.
2.4.2 Commercially available HME machines
There are several capable HME machines available from commercial suppliers.
The EV Group 520HE and the Suss SB6e and SB8e are adapted wafer bonding machines.
Jenoptik-Mikrotechnik offers three models of hot-embossing machines. Obducat has
three different embossing machines available. The capabilities of these machines are
26
summarized below. Obducat's product catalog gave only the maximum heat-up ramp,
rather than the time to heat from 60-180 0 C, and did not quote cooling performance. Suss
did not publish information of heat or cooling time in their web materials.
Supplier
Obducat [20]
JenoptikMikrotechnik
[21]
EV Group [22]
Suss MicroTec
[23]
Machine
NIL-2.5
NIL-4
NIL-6
HEX 01
HEX 02
HEX 03
520HE
SB 6e
SB 8e
Max
Max
Embossing Temperature
force (kN)
(C)
23
26
26
20
200
200
40
20
20
250
300
300
220
220
500
550
550
550
Heating
time
60-180 0 C
(min)
<1 0 C/s
<50 C/s
<50 C/s
7
7
7
6
?
?
Cooling Embossing
time
area
180-60 0 C
diameter
(min)
(mm)
65
?
102
?
152
?
130
7
130
7
120
7
200
5
150
?
200
?
Figure 2-9 Table of commercial hot embossing machines
Figure 2-11 Obducat NIL-4
Figure 2-10 EV Group 520HE
27
Figure 2-13 Suss SB6e
Figure 2-12 Jenoptik HEXOl
All of these machines are very able. All but the Obducat machines offer enclosed
embossing chambers permitting processing under vacuum, and many have built-in
automatic alignment systems for the tool and workpiece. All are able to control steadystate temperature to about ±1%. Most offer active cooling as an option.
2.5 The need for a new HME machine
While the many existing hot embossing machines are very capable, they have
certain deficiencies. The existing generation 1 machine lacks closed-loop control of
cooling and the ability to follow arbitrary temperature trajectories. Many commercial
machines have limited capability to follow arbitrary displacement or temperature
trajectories, and in most cases, direct control of the machine would be hidden behind a
layer of proprietary software and hardware. To successfully and completely address the
issues of manufacturing process control presented above, a new, custom-built machine is
needed.
28
The current work addresses this need for a new embossing machine in the MPCL.
The remainder of this thesis presents the design of this machine, covering the
development of design requirements, generation and selection of conceptual designs,
detailed design and analysis of predicted performance of the platen assembly and the
temperature control system.
29
CHAPTER
3
Goals for the new machine
3.1 Introduction
To extend research on manufacturing process control for hot micro-embossing, a
more capable machine is needed. This machine must enable the user to precisely control
the applied force and the displacement of the platens, and the temperature of the platens
and thus the workpiece and tool, and to control these parameters on arbitrary time
trajectories. The new machine should also incorporate improvements to allow waferscale processing and more automation.
The existing machine and most commercially available machines do not permit
investigation of higher heating and cooling rates, and thus the ultimate limits on
embossing cycle time are not yet known. A new machine must be capable of much faster
operation. Improvements in process automation over the generation 1 machine will
permit more precise control and repeatable experiments.
3.2 Probing spatial variation
The generation 1 embossing machine in the MPCL had a small maximum
workpiece size to ensure uniform distribution of heat and pressure across the workpiece,
thus eliminating a potential confounding factor. With an eye for increasingly complex
devices and higher production rates, the trend in embossing is unanimously towards
larger workpieces. Larger embossing areas permit larger devices or producing several
devices in one batch. Indeed, for a batch process such as this, increasing production rate
30
means either reducing cycle time-to which there is some physical limit-or increasing
batch size. Spatial variation of process parameters across the workpiece and tool is thus
an important disturbance factor needing investigation. Clever machine design can reduce
this variation, but it can never be totally eliminated. An effective strategy would also
involve tuning process parameters so that the sensitivity of the output to spatial variation
across the workpiece is minimized.
The new machine design should, of course, minimize the non-uniformity of
process parameters across the workpiece to the extent this is practical. The strategy of
the generation 1 machine was to reduce the size of the workpiece to the point that
variation was negligible; however, this precludes investigating the sensitivity of spatial
non-uniformity to process parameters. In order to probe spatial variation and to better
model potential industrial embossing, the new machine should accommodate a lager
workpiece.
Traditional photolithography has remained a convenient method for making
micro-features for embossing tools, so silicon wafers have long been the primary type of
embossing tool [24]. Several alternative tooling materials, such as etched glass,
electroplated metal, or laser-ablated silicon are produced with techniques designed for or
involving standard silicon wafers. Silicon wafers are available in standard diameters of
25, 50, 76.2, 100, 125, 150, 200, and 300mm. Wafer size has gradually increased over
several decades, with the smaller sizes mostly phased out, and the largest sizes only
recently introduced. The standard 100mm wafer, often referred to as a four inch wafer, is
the typical size used in the embossing literature, and most commercial embossing
machines are designed for this size. The 100mm wafer is still used by many university
31
and research centers because its equipment is smaller and less costly, and so is also a
typical platform for microfluidic, MEMS, and other related research.
Convenience, availability, and the consensus of the research community point to
100mm wafers as the target tooling size for the new HME machine. This size wafer
seems to be "just right" for this purpose, being large enough to model typical polymer
micro-devices, while still small enough that embossing forces and requirements for
uniformity of temperature and pressure are still reasonable.
3.3 Workpiece material and thermal requirements
PMMA will continue to be the target material, as it has favorable properties for
both fluidic and optical applications. The glass transition temperature for PMMA is
around 1 00 0 C, depending on molecular weight. This temperature is high enough for
room-temperature stability, and low enough that an embossing machine can easily heat
the workpiece above it. Prior work in the MPCL has used 1mm thick PMMA sheet at the
workpiece material, while much of the embossing literature has used thinner layers of
PMMA on silicon wafer substrates. The design of the new machine should not preclude
either of these uses. Above about 200'C, PMMA can be considered molten, so this
temperature forms the upper boundary of what should be considered embossing. Other
materials, such as polystyrene and polycarbonate, could also be embossed in this
temperature range.
3.4 Time-domain process parameters
A central goal for the new HME machine is that it be capable of probing the
effects of time-domain process parameters. These included embossing strain rate, hold
32
time, and heating and cooling rates. Non-linear force, displacement, and temperature
trajectories should also be possible. Even if the desired time-domain process parameter
trajectories are slow and linear, the capability for precise, fast, non-linear actuation will
perut robust rejection of disturbances.
The Instron model 5869 electromechanical load frame used for the existing
generation 1 HME machine is a very good force and displacement actuator. It is capable
of both force- and displacement-based control, and can produce a variety of waveforms in
either domain, as well as arbitrary user-programmed trajectories. The load frame is
capable of embossing forces up to 50kN. The Instron frame will be retained for the new
machine.
The existing generation 1 HME machine had relatively slow thermal response,
with heating time about 15min and cooling time about 5min. There will be an inevitable
interest in driving the cycle time for embossing to the minimum, so faster thermal
response will be needed in the new machine. Thermal response is primarily a function of
two aspects of the machine design-the available heat transfer power, and the thermal
mass of the components subject to thermal cycling. Thermal mass is here defined as the
product of an object's mass with is specific heat capacity, given in units of energy per
unit increase in temperature. Thermal response can therefore be measured in the simplest
sense by dividing the heat transfer power by the thermal mass, giving the maximum rate
of temperature change. Fast thermal response implies a combination of a powerful heat
transfer system and a low thermal mass. Thermal cycling should be limited to the fewest
components possible-that is only those in direct contact with the workpiece and tool.
33
These components should be designed with the thermal cycle in mind. They should
provide for uniform heat transfer to the workpiece and tool as well as adequate fixturing.
Ultimately, the thermal response is necessarily limited. A lower bound on what
should be expected of an embossing machine may be found by estimating the amount of
time a typical workpiece might take to reach thermal equilibrium with the machine. In
the expected embossing application for the new machine, a 1mm thick PMMA workpiece
will be cooled by contact on both sides, as shown in figure 3-1. The symmetry of this
situation can be invoked to simplify the model, as shown in Figure 3-2, where T is the
temperature distribution function, t is the time variable, x is the space variable, 1 is the
half-thickness of the workpiece, and T, is the temperature of the platen at the surface.
Platen
PMMA
'1 ..
.
Line of
symmetry
Platen
Figure 3-1 Thermal model of workpiece
x
aT0
Line of
symmetry
X~x=/
T(0,t)= T
Platen
Figure 3-2 Simplified thermal model of workpiece with boundary conditions
The time for the workpiece to reach thermal equilibrium with the platen can be
found by approximating the PMMA as an infinite slab. This approximation is valid
because the width of the workpiece is many times larger than its thickness. The one34
dimensional transient heat transfer equation is given in Equation 3-1, where T is the
temperature distribution, t is time, x is position, and a is the thermal diffusivity of
PMMA, defined as the ratio of thermal conductivity to the product of density and heat
capacity [25]. The position variable x has its origin at the bottom surface of the PMMA
and increases in the vertical direction.
82 T
ax
2
1 T
a at
Equation 3-1
The solution of this equation may be simplified by using the dimensionless
temperature (D, defined in Equation 3-2, where T is the temperature at a given point at a
given time, Ts is the temperature of the platen surface, and Tj is the initial temperature of
the PMMA. Substituting (D(x,t) into Equation 3-1 gives Equation 3-3.
0
T- T
Oi
T - T,
Equation 3-2
a(
a 2q
i
ax2
a at
Equation 3-3
The solution of this equation is assumed to be of the form shown in Equation 3-4,
where F is a function of only the position variable x, and G is a function of only the time
variable t. Substituting this solution into Equation 3-3 gives Equation 3-5. Dividing the
equation by the product FG gives Equation 3-6.
1(x,t)= F(x)G(t)
Equation 3-4
G
d2 F
1
d2 G
=-F
dt 2
a
dx2
2
35
Equation 3-5
I
d2 F
-
Fdx2
=-
1
dG
a
dt
Equation 3-6
The right and left sides of Equation 3-6 are independent, so they must be equal to
the same constant. This equation can be separated and re-arranged to give Equation 3-7
and Equation 3-8, where -C 2 is the separation constant. These equations may be easily
solved to Equation 3-9 and Equation 3-10 respectively. These solutions illustrate the
reason for choosing -C 2 . The separation constant is chosen to be negative because the
temperature difference, and thus the dimensionless temperature, is expected to decay to
zero over time as the PMMA comes into equilibrium with the platen. The constant is
squared to make the determination of the sine and cosine coefficients cleaner.
dG =-C 2
aG
dt
Equation 3-7
d2
F+C2F
2
=0
dx
Equation 3-8
G(t) = ec
2
at
Equation 3-9
F(x) =
{sin(Cx)
cos(Cx)I
Equation 3-10
The lower surface of the PMMA is subject to a constant temperature, and the
upper boundary is constrained to have zero heat transfer through the boundary. These
constraints give the Dirichlet condition in Equation 3-11, and the Neumann condition in
Equation 3-12, where 1 is the half-thickness of the PMMA.
36
Equation 3-11
ax x=1
=0
Equation 3-12
Because the value of cos(Cx) is 1 at x=0 while the value of ID(O,t) is constrained
to a constant value of zero, the cosine solutions can be discarded. The basic solution to
Equation 3-3 is thus given by Equation 3-13. The value of C can be found by imposing
the Neumann condition, giving Equation 3-14, where n is a positive integer.
(=
e-C "' sin(Cx)
Equation 3-13
2 "t cos(Cl)
0 = e-c
C (2n -1)
2
r
1
Equation 3-14
The full solution is the sum of all the particular solutions. This summation is
given in Equation 3-15 where Bn is the nth coefficient of the Fourier sine series given by
Equation 3-16 where (o is the temperature distribution at t=0 [26]. This distribution is a
constant equal to 1 for x>0 because the temperature throughout the PMMA is equal to the
initial temperature. The value of the function for x<O is meaningless, so the Fourier
integral is only evaluated for positive x.
i(Dx,t)= IB,e
-c" sin(Cx)
n=1
Equation 3-15
B
=
1
(-Do
sin(nx)dx
rc
-1
1)
(cos(nEr)qnir
Equation 3-16
37
The summation solution may be approximated by its first term, where n=1, giving
Equation 3-17. The temperature at the center of the PMMA is found by setting x=l,
leaving a purely exponential function. The time constant of this function is given by
Equation 3-18 [27].
(D(x, t)= 1 e(/2)"' sin - x
Equation 3-17
21 )2
1
7r
a
Equation 3-18
Using the standard definition the settling time as four times the time constant, the
half-thickness of 0.5mm, and the diffusivity of PMMA of 1.23x10- m 2 /s, the settling
time for the temperature at the midline of the workpiece is found to be 3.29 s [27]. This
time sets a lower boundary on the thermal response time of the machine. In practice, the
heating or cooling time for the embossing machine will be longer, since the platens and
heat transfer system will themselves have thermal masses. Without knowing the impact
of fast heating and cooling on quality of the final part, it is difficult to say exactly how
fast is "fast enough." There will be limits to how small the platens can be made while
still accommodating the heat transfer system and allowing for fixturing of the workpiece
and tool, and similarly there is a practical limit on the power of the heat transfer system.
These limits are not fixed, but are based on feasibility, complexity, and cost.
3.5 Automation
Automation of the HME machine is beneficial in that it increases the repeatability
of experiments vs. manual control, and it makes experiments more convenient for the
38
user. The Instron load frame permits automated execution of programmed force and
displacement waveforms. In the generation 1 machine, the heaters are controlled to a
given set point automatically, but this temperature must be adjusted manually. Cooling is
entirely manual by opening and closing water valves. The temperature control and heat
transfer system for the new machine should be completely automated, enabling the
machine to follow pre-programmed temperature profiles just as it does for force and
displacement.
3.6 Workpiece and tool fixturing
In the original generation 1 HME machine, problems were encountered in
fixturing the workpiece and tool. The PMMA workpiece was clamped to the bottom
platen around its periphery using a copper plate with a hole in the center through which
the tool could pass. Embossing tools were mounted to the upper platen. In order to pass
through the hold on the clamp plate, the tool had to be aligned properly, so the mounting
holes on the tool fixture were oversized, allowing it to move a small distance laterally.
The procedure involved leaving the tool fixture screws loose, running the platens together
until the tool fixture mated with the hole in the clamp plate, then separating the platens
and tightening the tool fixture screws. In practice, this procedure proved difficult
because the tool fixture would sometimes move before the screws were tightened, and the
process had to be repeated. Misalignment between the tool fixture and the workpiece
clamp plate caused damage to the tool fixture on one occasion. The new machine should
incorporate improvements in the workpiece and tool fixturing that will remove the
necessity that the fixtures be re-aligned after every tool change.
39
Both silicon wafers and machined copper pieces have been used in the generation
1 machine. Fixturing copper tools presented little difficulty because they could have
integral holes for screws. Because silicon wafers are quite thin and brittle, they proved
difficult to mount. Perimeter clamping was not possible because the clamp would intrude
into the PMMA workpiece. Instead, silicon tools were affixed to a tool post using hightemperature epoxy. This method proved unreliable because it was difficult to ensure that
the silicon tool was parallel to the tool post surface. Voids in the epoxy were sometimes
present as well. Because of misalignment and possibly voids, silicon tools would
sometimes fracture under embossing forces. Adhesion between the epoxy and the silicon
tools was poor, so they would sometimes break away from the tool post and remain
embedded in the PMMA workpiece. The new machine should incorporate a more
accurate and reliable method for fixing thin tools such as silicon or glass wafers.
3.7 Project scope
The current work addresses the development of concepts and the detailed design
and analysis of the platen assembly and the temperature control system for the new HME
machine. It does not include developing an integrated control program for both the load
frame and the heat transfer system. The planned design does not include an enclosed
chamber for embossing under vacuum, nor does it include an active system to ensure the
platens are precisely aligned.
Roos et al compared embossing under vacuum to embossing at ambient pressure,
and found that uniformity across the workpiece was greater under vacuum [14]. They
used a tool with features only 280nm tall, and PMMA only 300nm thick, so it is not clear
that their results apply to features several microns tall and PMMA up to 1mm thick.
40
Bacon et al [15] using the same embossing apparatus found that uniformity across the
part also varied with other processing conditions such as embossing temperature and
force, suggesting that with proper selection of embossing conditions, vacuum may not be
necessary for uniform imprinting. Indeed, results by Ganesan [17] and continuing
experience with HME in the MPCL have demonstrated very good embossing for features
at this scale, without the need for vacuum. A vacuum chamber would increase the
complexity of the embossing machine, while the necessity of a vacuum environment has
not yet been demonstrated.
The current design also does not incorporate an alignment system for the two
platens. Such a system could be anything from passive alignment using beams and slide
bearings to closed-loop, actuated control of the platen alignment in up to five degrees of
freedom. Past experience in this lab has shown that the alignment accuracy of the Instron
load frame is adequate to produce good results. Any misalignment resulting from
tolerances in the machined parts does not vary over time, so these misalignments can be
compensated for with shims and other adjustments. Although the current design does not
include a vacuum chamber or active alignment system, it should not preclude their
addition at a later date.
3.8 Summary of goals for the generation 2 HME machine
The generation 2 HME machine is intended to meet the demands of an extensive
investigation of HME from a manufacturing process control perspective. It is intended to
enable experiments across the practical range for all process parameters, including time
domain parameters. It will accommodate a workpiece up to 100mm in diameter to study
spatial variation across the part, it will permit faster experiments and probing of the
41
ultimate limits to HME cycle time, and it will enable full automation of both the force
and displacement control and the temperature control. The new machine should also
permit more reliable fixturing for tools and workpieces.
42
CHAPTER
4
Concept development and evaluation
4.1 Introduction
The generation 2 HME machine will incorporate numerous improvements over
the existing machine. The most important of these improvements addresses the lack of
precise control over the temperature-time trajectory. Reducing the cycle time,
incorporating full automation, and improving fixturing are also important requirements
for the new machine. To proceed with the design for the new machine, conceptual
designs to meet each important function were developed and evaluated. The concept that
proved most likely to meet the requirement was selected for the detailed final design.
4.2 Temperature control
Improving control over the temperature-time trajectory and reducing heating and
cooling time are the central, defining requirements for the new machine design. The
method for heating and cooling the platens to control the temperature of the workpiece
will drive the design of the platens themselves, so the temperature control subsystem
must be defined before the design of the platens can proceed.
The temperature control system must be able to add or remove heat energy to the
platens in order to change their temperatures. There are three domains of heat transferconduction, convection, and radiation-and all should be considered for the temperature
control system. Conduction will necessarily be the operative mode of heat transfer within
the platen and fixture assembly, but the manner of heat transfer to and form this assembly
43
could be of another form. A source of heat energy is needed to add energy to the platen
assembly and increase its temperature, and an energy sink is needed to remove energy
and decrease temperature. Many methods for generating and dissipating heat energy
exist, however some are better suited to a laboratory environment.
Whatever method is chosen must be capable of producing heat transfer rates
sufficient to meet the heating and cooling time goals. Rough estimates for the minimum
total power may be found by multiplying the thermal mass of the workpiece and tool by
the desired temperature change, and dividing by the desired time to change. The
workpiece and tool are taken to be 100mm diameter circles. The workpiece is PMMA
1mm thick, and the tool is silicon 0.5mm thick, giving thermal masses of 13.6 J/0 K and
6.4 J/0 K respectively. For a temperature change from 25*C to 150'C, this gives a total
energy change of 2.5kJ. To accomplish this temperature change in 2 minutes would
require an average power of 21W. In practice, however, some of the heat input must go
to changing the temperature of the platens, so the required power will be higher.
4.2.1 Thermoelectric (Peltier)
A fully electric system would have the advantage of being physically simple and
clean. Furthermore, the mechanism of heat transfer would be in the same domain as the
electrical control and feedback signals, reducing the number of conversions between
domains of energy and resulting in higher efficiency. This is made possible via the
Peltier effect, whereby a temperature difference is created by current passing through
dissimilar materials. Because the polarity of current determines the direction of heat
transfer, heating and cooling may be accomplished using the same hardware simply by
reversing the current. Peltier effect heaters and coolers are available with power ratings
44
sufficient for the heating and cooling time goals, and some can operate at typical
embossing temperatures [28].
Although Peltier effect heating and cooling is attractive for temperature control in
embossing, certain physical realities render it infeasible. The first problem with solidstate Peltier effect devices is structural. Commercially available thermoelectric devices
are not designed to resist compression or tension loads, so some additional structure
would be needed to support the workpiece and tool. This structure would increase the
total thermal mass, so power requirements would increase. The second problem is
thermodynamic. The Peltier effect does not create or destroy heat energy, but merely
moves it from one place to another. Thus, when heating the workpiece and tool, the other
junction must have a source of heat, and when cooling, this junction must be exposed to a
heat sink. To produce the large changes in temperature encountered in embossing would
still require an additional heat transfer system to alternately heat and cool the Peltier
devices themselves, although the requirements on this system would not be as stringent as
on one that heated and cooled the workpiece and tool directly.
The Peltier effect has been successfully exploited for cooling of microelectronic
devices, but is not adequate for the bulk heating and cooling encountered in hot
embossing. Peltier devices could conceivably be integrated into the platens to produce
relatively small, localized temperature variations.
4.2.2 Mixed electric & fluid
The method of temperature control employed by the generation 1 machine
consists of a mixture of heat transfer strategies. Heat energy is added to the platens by
conversion from electrical energy though Joule heating. The heat flux from electric
45
cartridge heaters is controlled by adjusting the current flowing through them. Heat is
removed from the platens by convection to water flowing through passages in the platens.
Precise temperature control along arbitrary trajectories is not possible with the
original setup, but this method of temperature control could be adapted for the new
machine to allow this capability. The flow rate of cooling fluid could be adjusted to
modulate cooling while the current through the heaters is adjusted to control heating.
Balancing the heat flux into the platens from the electric heaters with the flux out of the
platens to the cooling fluid would permit temperature control. In practice, this strategy
would be difficult to implement. The relationship between flow rate and heat flux for
convection in tubes is extremely non-linear, exhibiting an abrupt jump in heat transfer
across the transition from laminar to turbulent flow.
This effect could be mitigated by running the cooling fluid at a constant flow rate
and varying the output of the electric heaters. For cooling, the heaters would be adjusted
so the heat flux into the cooling fluid is greater than the heat generated by the heaters, and
for heating the current would be increased so the heaters overpower the convective
cooling. This strategy requires that the heaters be very powerful, effectively doubling the
power needed. Temperature control would be inherently difficult because one has
control over heat fluxes, rather than temperature directly. A steady temperature is
accomplished by making the heat flux from the heater and the heat flux into the cooling
fluid equal, and changes in temperature are effected by adjusting the net heat flux into the
platens. When the heat fluxes are balanced, temperature will not change, but this steady
temperature does not depend on the values of the heat fluxes, indeed any absolute heat
flux could maintain any steady temperature so long as it is balanced.
46
Even if these control problems are solved, there still exists a physical drawback to
the mixed electric-fluid strategy. The heat source and sink are physically separated in the
platens, so there will be intense thermal gradients between the regions near the very hot
heater and the regions near the very cold cooling channels. The platens would have to be
designed very carefully to make sure these gradients to not impinge on the workpiece or
tool, ensuring uniform heating and cooling.
4.2.3 All fluid
An alternative strategy combines the functions of heat source and sink into a
fluid-based temperature control system. In this concept, fluid flows continuously through
passages in the platens, and the temperature of this fluid is controlled by a separate
system. This strategy removes the ultimate heat source and heat sink from the platen
assembly, simplifying its design and potentially reducing its thermal mass as well. The
temperature of the platens can be controlled directly by controlling the temperature of the
circulating fluid. A temperature control problem still exists of course, but one is freer in
solving it without the constraint that the control system be contained in the platen
assembly. The simplicity of this design with respect to the platen assembly and the
control problem recommends it strongly.
4.3 Concepts for thermal fluid supply system
The selection of an all-fluid temperature control strategy greatly simplifies the
design of the platens, but it has transplanted to control design into a separate system.
Several strategies for fluid temperature control exist.
47
4.3.1 Bulk heating & cooling
The simplest and most often applied method for controlling the temperature of a
circulating fluid is to heat and cool the bulk of the fluid as needed. The fluid is forced
through a heat source and heat sink, and the source and sink are activated depending on
whether the temperature of the fluid should be raised or lowered. This strategy has the
advantage that the fluid circulation system is physically simple, with no branches or
control valves needed. Unfortunately, it can be very slow. When heating the fluid, the
thermal mass of the heat sink is included, while when cooling, the heater must be cooled
as well. The entire bulk of the fluid must also be heated and cooled.
Platens
.......
Heat
Sink
Pump
Heat
Source
Figure 4-1 Bulk heating & cooling of fluid
4.3.2 Separated streams
This problem can be mitigated by separating the fluid into hot and cold streams.
Fluid can be diverted to either a heater or cooler as needed. When actively heating, flow
through the cold branch virtually ceases, so the heat sink can remain cold. Similarly, the
heat source retains its heat while the system is cooling the fluid. To maintain
intermediate temperatures, the hot and cold streams can be mixed in the proper ratio.
48
Indeed the exact temperature of the fluid leaving the heater and cooler need not be
controlled very precisely, as the control valves determine the final mix of fluid from the
two branches. The valves can be controlled via temperature feedback from their outlets,
so any disturbance in the temperature of the hot or cold branch could be rejected quickly
by changing the mixing ratio. The temperature of the fluid can then be controlled to any
temperature within the band between the hot and cold branch temperatures.
Platens
Heat
Sorc
Source
Control
Valve
P
Heat
Sink
Figure 4-2 Separated hot & cold streams
There are many options for the heat source and sink. The most convenient heat
source for a laboratory application is Joule heating by electric current. A wide variety of
electric heaters for fluid systems is commercially available. Typically, these take the
form of a cluster of electric heating elements within an enclosed chamber that is ported
with an inlet and outlet. Fluid gains heat as it flows around the heating elements as it
passes through the chamber. An example of this type of heater is shown in Figure 4-3,
courtesy of Process News Magazine.
49
Figure 4-3 Photo of an electric circulation heater [29]
For cooling, a heat exchanger will transfer heat from the circulating fluid to some
other medium, typically another fluid. Candidate heat sinks are the atmosphere, water, or
refrigerant. Air is limited in its ability to absorb heat, so a large heat exchanger would be
needed. Using the ambient air as the heat sink would also tend to heat the room in which
the machine runs. Cooling with refrigerant requires a second circulation system with its
own pump and heat sink. This method of cooling would be overly complex and
redundant. The best option would be to use a continuous flow of tap water to cool the
circulating heat transfer fluid. The oil and water would flow in opposite directions
through a heat exchanger. A plate-and-frame heat exchanger consists of a stack of thin
plates closed at the edges. The plates are connected in such a way that the two fluids
flow through alternating layers, as shown in Figure 4-4, courtesy of Tranter PHE. The
50
large surface area and narrow gaps between the plates make this type of heat exchanger
very compact and effective. Several sizes and shapes are commercially available.
END
14EAT TRANSFER
PLATES
COLD
PLAATE
END
HOT
Figure 4-4 Diagram of plate and frame heat exchanger [30]
4.4 Selection of working fluid
With the separated-stream all-fluid temperature control concept selected, the
working fluid for the heat transfer system should be selected. The fluid should have good
thermal properties, such as a low specific heat and high thermal conductivity. Liquids are
thus favored over gases. This fluid must tolerate the temperature range encountered in
embossing PMMA, from ambient temperature up to 200'C. To minimize the complexity
of the fluid heat transfer system, the system should not need to be pressurized. Thus, the
heat transfer fluid should not boil in this temperature range. This requirement excludes
pure water as well as aqueous solutions of ethylene glycol.
There are many other available heat transfer oils such as hydrocarbon and siliconbased oils. It is important that the working fluid be convenient for a laboratory
51
environment, and so should not require careful handling or an inert atmosphere within the
system. Aromatic hydrocarbon oils are thus excluded because of their toxic properties.
The working fluid should have a relatively low viscosity across the operational
temperature range to minimize the required pump power as well as the pressure losses
through the system.
After an extensive review of commercially available heat transfer oils conducted
by Grant Shoji (an S.M. student at MIT and colleague of the author), Paratherm MR was
selected. This fluid is a paraffinic hydrocarbon oil. It has favorable properties, with a
room temperature viscosity about 4 cP, comparatively high thermal conductivity, and a
boiling point above 300'C. It is non-toxic and nearly odorless, and has handling
requirements similar to typical lubricating oils. As with all heat transfer oils, an inert
atmosphere is preferable, however the fluid is thermally and chemically stable enough
that this measure is not required. Instead, a cold trap can be used. The cold trap isolates
the working fluid from the atmosphere, just as a drain trap prevents sewer gas from
entering a household plumbing system (see section 6.7). Properties for Paratherm MR at
many temperatures are tabulated in appendix A. 1.
4.5 Workpiece and tool fixturing
The new machine must improve the fixturing for the workpiece and tool. The
new design should eliminate the need to re-align the fixtures after every tool change. The
fixture for wafer-based tools must be sturdy and reliable. The fixtures must support
tensile loads so automatic de-embossing is possible.
Perimeter clamping as used on the generation 1 machine is simple and reliable,
but because the workpiece and tool must be pressed together, only one or the other can be
52
clamped in this fashion. The perimeter clamps for the workpiece and tool could be made
with interdigitated gaps so they will mesh together, but some clamp material would still
impinge on the workpiece. This would also tend to concentrate contact forces in small
regions on the tool. Because silicon is fairly brittle, stress concentrations should be
avoided.
Vacuum chucks are commonly used to fix silicon wafers in traditional microfabrication machines. These chucks have several small holes or grooves in their surfaces,
and these are connected to a vacuum pump. The wafer is placed over the holes or
grooves and the pressure of the atmosphere against the wafer holds it in place. The
clamping force is limited to the projected area of the holes and grooves multiplied by the
ambient pressure. For a 100mm wafer, the maximum possible atmospheric clamping
force is 796 N. A projected area of 100% is clearly impossible for a vacuum chuck, but a
50% area may be more feasible. A total vacuum is also not feasible. For a vacuum
chuck with 50% projected area and a vacuum of 1 kPa absolute pressure, the clamping
force would be 394 N. The clamping force for the vacuum chuck represents the upper
limit on the de-embossing force the machine can exert. Anti-adhesion layers may be
applied to the tool to reduce de-embossing forces.
4.6 Summary of conceptual design
The generation 2 machine will have platens heated and cooled by Paratherm MR
flowing through internal passages. The temperature of the fluid will be controlled by
diverting the flow through either a heat source or sink, or by mixing streams of fluid from
each branch. The displacement of the platens and the embossing force trajectory will be
controlled using the existing Instron load frame. The PMMA workpiece will be clamped
53
at its perimeter as in the generation 1 machine, while the tool will be held in a vacuum
chuck. The new machine should be capable of heating from ambient to embossing
temperatures in about 2 min, and cooling time should also be about 2 min. The precise
temperature trajectory in both heating and cooling should be controllable. These
improvements in function and control will permit extensive experiments to study hot
micro-embossing as a manufacturing process.
54
CHAPTER
5
Design of the platen assembly
5.1 Introduction
The platen assembly is the most critical portion of the embossing machine. The
platens and fixtures are in contact with the workpiece and tool, and so have a direct effect
on quality. The platens and their internal passages must provide for even distribution of
heat transfer. The fluid manifolds must distribute even fluid flow among the tubes in the
platens. The platen assembly also includes the interface between the load frame and the
actual platens and must maintain alignment between the workpiece and tool and support
them under embossing and de-embossing loads. The platen assembly must also isolate
the components that are thermally cycled both to reduce the thermal mass and to protect
sensitive components of the load frame from heat. A basic schematic of the platen
assembly including the platens and components for fixturing, mounting, and insulation is
shown in Figure 5-1. The detailed design for each of these components is discussed in
the following sections.
55
Mount to Instron crosshead
Insulation
Platen
Maifo d
Manifold
Too fixture
Workpiece fixture
Platen
Matifold
Manifold
Insulation
Mount to Instron frame
Figure 5-1 Basic schematic of platen assembly
5.2 Design of the platens
The platens serve two important functions: they support the workpiece and tool
under embossing loads, and they transfer thermal energy to and from the workpiece and
tool. The platens themselves are part of the thermal circuit, so their thermal mass must be
considered. Indeed, the thermal mass of the platens will dominate the temperature cycle
because the workpiece and tool are quite small. In the generation 1 machine, the plates
were massive blocks of copper in order to even out the thermal flux from concentrated
heat sources. In order to minimize the thermal mass of the platens, the source of thermal
flux should itself be even. The minimal geometry consists of thin platens with several
small internal passages for thermal fluid. In this configuration, the platens are merely an
intermediary between the thermal fluid and the workpiece, serving to average out the
thermal flux between the passages and to provide structural support for the workpiece.
Pure copper is the ideal material for this application because of its high thermal
conductivity.
56
Figure 5-2 Minimal platen design
5.2.1 Parametric model of platen performance
The effectiveness of convective heat transfer between the thermal fluid and the
platens will determine their thermal performance. Convective heat transfer is governed
by Newton's law of cooling, where q is the heat transfer rate, A is the surface area
exposed to convection, h is the average convection coefficient across this area, T, is the
temperature of the surface, and Tm is the mean temperature of the fluid [25].
q = hA(T - Tm)
Equation 5-1
The average heat transfer coefficient depends on the conditions of fluid flow
inside the tube. The most important consideration is whether the flow is in the laminar or
turbulent regimes, as determined by the Reynolds number. The critical Reynolds number
for transition to turbulence in internal flow is 2300 [25]. The heat transfer coefficient is
customarily calculated from the dimensionless Nusselt number, which is defined below,
where h is the convection coefficient, L is the length of the tube and k is the thermal
conductivity of the fluid [25].
Nu = hL
k
Equation 5-2
For laminar flow through circular tubes with uniform wall temperature, the
Nusselt number is a constant equal to about 3.657 [31]. For turbulent and transitional
flow, the situation is much more complex. Nusselt numbers are found using empirical
correlations, of which there exists a great variety. One of the more versatile correlations
57
is attributed to Gnielinski, and is valid for transitional and turbulent flows with Reynolds
number between 2300 and 5x 106 and Prandtl numbers between 0.5 and 106 [32]. This
correlation is given in Equation 5-3. The friction factor f is given by Equation 5-4 for
turbulent flow in smooth circular tubes, and is valid for the same ranges of Reynolds and
Prandtl number.
N_
1000)Pr
_(f/2)(Re-
1+12.7(f/2)
(Pr213 _
Equation 5-3
f
= (1.58 ln(Re)- 3.28) 2
Equation 5-4
These correlations were implemented in a MatLab script in order to show the
relationship between the heat transfer coefficient and the tube diameter and flow velocity.
The MatLab script contains a parametric model of the platens based on the minimal
design shown in Figure 5-2. The model platen has a thickness of three times the tube
diameter. The width is determined parametrically. The minimum tube spacing is one
tube diameter, so the minimum platen width of 100mm is divided by twice the tube
diameter and rounded to the largest integer, giving the minimum number of tubes that
will span the minimum platen width. Two tubes are added to mitigate edge effects
(discussed in section 5.4.2). The final width is twice the number of tubes times the tube
diameter, plus one diameter to wall off the last tube. The MatLab script code for finding
the number of tubes and the width of the platen is given below, with D representing the
tube diameter. All units are SI. The length of the platen-and thus the tubes-is taken to
be equal to the width. The full MatLab code is available in appendix C.2
%Calculate platen characteristics based on tube diameter D
nchan=ceil(.l/(2*D))+2; %number of tubes per platen
58
L=nchan*D*2+D; %Length of platen (equal to width)
The script evaluates the convection correlation for different combinations of tube
diameter and flow velocity. The properties of Paratherm MR vary widely with
temperature. For this analysis, the fluid is considered at room temperature, or 25 0C,
because this represents the lowest operating temperature and the "worst case"
performance of the fluid. As temperature increases viscosity decreases, so the Reynolds
number increases strongly. Because the Nusselt number strongly depends on the
Reynolds number, the convection coefficient will be much higher at higher temperatures.
The values of the convection coefficient found in this manner are shown in Figure 5-3.
The convection coefficient is seen to be strongly dependent on flow velocity. A large
increase in the value of the coefficient is evident at the transition from laminar to
turbulent flow.
59
2'7000-E6000
5000
4000
8 3000
10
15
10
5
Tube diameter (mm)
0
0
Flow velocity (m/s)
Figure 5-3 Convection coefficient results from parametric model
The convection coefficient is somewhat dependent on tube diameter. The
diameter of the tube is even more important in determining the overall geometry of the
platen, and thus the thermal mass and time-domain performance. This relationship is
shown in Figure 5-4. The line is angular because of the discrete rounding involved in
finding the number of channels.
60
8
76
:S3-2-
0
2
4
6
8
Tube diameter (mm)
10
12
Figure 5-4 Dependence of platen mass on tube diameter
From the above figures, it is evident that the ideal platen will have tubes as small
as possible and flow velocity as high as possible. There are two countervailing trends
that place limits on the smallest tube diameters and largest flow velocities. The first of
these is the pressure drop associated with forcing fluid through small tubes at high
velocities. Pressure drop in a circular tube is given by Equation 5-5, where Ap is the
pressure drop, f is the friction factor, L is the length of the tube, D is its diameter, p is the
fluid density, and V is the flow velocity [33].
fL pV
D
2
2
Equation 5-5
61
The same MatLab script was used to calculate values of this pressure drop for the
same combinations of diameter and flow velocity. Again, room temperature fluid
properties were used because this is the "worst case" for pressure drop. The values of
pressure drop computed from the model are shown in Figure 5-5.
80Cz
600
40V
C,,
20
0
15
10
-
10
55
dimeter (mm)
TubhPerim
T"h
0
0
Flow velocity (m/s)
(Mfiz)
Figure 5-5 Pressure drop from parametric model
The pressure drop is somewhat dependent on flow velocity, and increases sharply
as tube diameter decreases. Discontinuities at the transition from laminar flow to
turbulent flow are also visible. A high pressure drop across the platens is undesirable
because it increases the pressure within every component upstream of the platens. The
62
power of the pump must also be greater to overcome the higher pressure loss in the
system. High flow velocity is also detrimental. A higher flow rate will cause greater
pressure drops in other components in addition to the platens, and will require a larger
pump and more powerful heater. The computed values for volume flow rate are shown
below.
400
E
1-00300
200
E 10001
15
Tube diameter (mm)
0 0
Flow velocity (m/s)
Figure 5-6 Volume flow rate from parametric model
The results from the parametric model give guidance on selecting the tube
diameter and flow velocity for the generation 2 HME machine. In order to minimize the
mass of the system as well as the required fluid flow rate, the tube diameter should be as
63
small as possible, but not so small as to cause a severe pressure drop. The flow velocity
should be chosen to maximize heat transfer performance while keeping pressure drop
low.
5.2.2 Manufacturability & selection of tube diameter
It has been found that the tube diameter should be as small as possible. The range
of possibility is defined by the processes available to produce long, narrow holes in
metals. Laser drilling and water-jet drilling can produce deep holes with small diameters
holes, but these holes are slightly conical. Gun drilling is capable of producing long,
straight holes in a variety of materials, and drill bits are available in lengths of several
inches with diameters as low as one eighth of an inch. Another method for producing
small diameter holes would be to mill small channels in the face of a copper plate and
then braze another copper plate over them. Capillary action would cause molten brazing
alloy to fill very small channels, so a practical limit on channel width is again roughly
one eighth of an inch. For this brazing process, there will be two plates with a total of
twelve machined surfaces, four of which-the tops and bottoms-have critical
tolerances. Brazing the two plates together would also require alignment holes and pins,
bringing the total number of critical features to ten features among four parts. For a gundrilled platen, there is only one part with six machined surfaces, and only the top and
bottom have critical tolerances. Both processes must be performed by vendors outside
MIT, and they have comparable cost. Gun drilling reduces manufacturing complexity, so
it is the preferred method for producing the fluid tubes in the platens.
64
5.2.3 Specifying a flow rate
For a tube diameter of one eighth of an inch, or 3.175 mm, the parametric
minimal platen design would have 18 tubes and an overall length of 4.625 in, or 117.475
mm as found by the MatLab script discussed in section 5.2.1. With the tube diameter set,
the flow velocity can now be chosen. From Figure 5-3 and Figure 5-5, it is evident that
the pressure drop is reduced and the heat transfer effectiveness is increased when the
fluid flow is in the turbulent regime. A minimum flow velocity can thus be found from
the critical Reynolds number. Using Equation 5-6, where V is the flow velocity, D is the
tube diameter, p is the fluid density, and pt is the fluid viscosity, this minimum velocity is
found to be 4.33 m/s. For two platens with 18 tubes each, this would require a total flow
rate of 19.5 gpm, or 1.23x10-3 m 3/s.
Re = VDp -2300
Equation 5-6
The results of the parametric model discussed in sub-section 5.2.1 show that
increasing the flow velocity will increase the heat transfer rate, with the penalty of
increasing pressure drop across the platen and the total flow rate. The relationship
between convection coefficient and flow velocity is shown in Figure 5-7, and pressure
drop is shown in Figure 5-8.
65
6000
25000 -
E
S4000 U)
E 30000
020000
0 1000-
0
4
6
8
10
Flow velocity (m/s)
12
14
Figure 5-7 Convection coefficient vs. flow velocity
The discontinuity visible at left occurs at the transition to the turbulent convection model for flow
velocity above 4.3 m/s
66
25
200L
15 0
10
5--
0
4
6
10
8
Flow velocity (m/s)
12
14
Figure 5-8 Pressure drop vs. flow velocity
The discontinuity visible at left occurs at the transition to the turbulent convection model for flow
velocity above 4.3 m/s
From these graphs alone, it is difficult to determine what flow velocity will
produce the desired heat transfer performance at the minimum pressure drop and flow
rate. The parametric model can be used to estimate the time-domain performance of the
platens to give better insight into selecting the proper flow velocity. The lumped
capacitance method can be used to estimate the transient thermal behavior of the model
platen. The lumped capacitance method assumes that the resistance to heat transfer
within a solid is low compared to the resistance to heat transfer between the solid and a
67
convective fluid [25]. The validity of this assumption is tested by calculating the ratio of
external heat transfer to internal heat transfer by Equation 5-7, where h is the convection
coefficient, V is the volume of the platen, k is the thermal conductivity of copper, and A
is the surface area exposed to convection. This ratio is known as the Biot number, and it
varies with h, which varies with the flow velocity as shown in Figure 5-9. The lumped
capacitance assumption is considered valid for Bi<O. 1, as is the case for this range of
flow velocity.
Bi =hV
kA
Equation 5-7
0.1
0.081
-i)
E
0.061
tO 0.041
0.02
It
4
I
6
I
I
8
10
Flow velocity (m/s)
I
12
14
Figure 5-9 Biot number vs. flow velocity
The discontinuity visible at left occurs at the transition to the turbulent convection model for flow
velocity above 4.3 m/s
68
Using the lumped capacitance method, transient heating and cooling of the platen
can be represented by a first-order differential equation. In Equation 5-8, h is the
convection coefficient, A is the surface area subject to convection, Tp is the temperature
of the platen, Tm is the fluid temperature, p is the density of copper, V is the volume of
the platen, Cp is the specific heat capacity of copper, and t is time. Substituting O=Tp-Tm,
the differential equation can be solved to Equation 5-9.
- hA(T - T) =p
C, dT
dt
Equation 5-8
hA
O
t_
e -= t
--
Oi
Equation 5-9
The quantity hA/pV-C, is the inverse of the time constant T [27]. The time
constant is a convenient measure of the transient thermal behavior of the platen, and can
be computed using the parametric model.
69
12
10-
4-
4
6
8
10
Flow velocity (m/s)
12
14
Figure 5-10 Time constant vs. flow velocity
For each incremental increase in flow velocity, there is a diminishing return on
decreasing the time constant, while pressure drop and total flow rate continue to increase.
Higher pressure drops and flow rates will require more powerful pumps and heat
exchangers. The final total design flow rate was chosen as 40 gpm through two platens,
corresponding to a flow velocity of 8.85 m/s through the tubes in the platens, and a
pressure drop of about 11.7 kPa. The convection coefficient for 25'C fluid at this flow
velocity is about 3900 W/m 2 K, giving a time constant of about 4.8 s. For comparison, a
flow rate of 30 gpm would reduce pressure drop to about 6 kPa, but the time constant
would be about 8 s. Increasing the flow rate to 50 gpm would only reduce the time
constant to 3.8 s. The final choice is a subjective balance between the benefit of a lower
time constant and the costs of increasing system complexity.
70
5.2.4 Final platen design
The final platen design incorporates holes for mounting and a recessed area to
mate with the manifolds. The top and bottom platens are identical. The larger holes are
for the screws that will connect the platen to the carrier plate, and the smaller holes will
be tapped for screws to hold the workpiece clam and vacuum chuck.
Figure 5-11 Final platen design
5.3 Fixturing and mounting
The two main functions of the platen assembly are to transfer heat into and out of
the workpiece and tool, and to support them under embossing and de-embossing loads
while maintaining alignment. Various improvements over the generation 1 machine are
to be incorporated into the design of the new machine. The fixture should eliminate the
71
need to re-align the workpiece clamp and the tool chuck. A vacuum chuck will hold the
tool, and a perimeter clamp will hold the workpiece.
5.3.1 Mounting the platens
In the generation 1 machine, it was necessary to re-align the tool chuck and
workpiece clamp after a tool change. This was because alignment was adjusted using the
mounting screws for the tool holder. In order to eliminate this problem, the new vacuum
chuck and workpiece clamp plate will not be adjustable relative to the platens. To
provide for adjustment of the lateral alignment of the platens, one of them must be
movable. This will be accomplished by incorporating slotted mounts where the bottom
platen attaches to the Instron load frame.
A T-slotted table will be added to the load frame, as shown in Figure 2-1. The
four holes correspond to holes on the load frame, and the T-slots will allow lateral
adjustment of the bottom platen. The bottom platen will be mounted to a carrier plate
which has holes slotted from front to back as in Figure 5-13 Bottom carrier plate. These
slots will allow the bottom platen to be adjusted from front to back.
72
Figure 5-12 T-slotted table
Figure 5-13 Bottom carrier plate
With the bottom platen adjustable in both x and y, the top platen need not be
adjustable. The top platen will be mounted to a carrier plate, which will in turn be
mounted to the top compression anvil. The platens will be connected to the carrier plates
73
by four screws fitting tapped holes in the plates. The carrier plates will be made of plain
carbon steel for strength and durability.
Figure 5-14 Top carrier plate mounted to anvil
The Instron anvil has a special clevis pin connection for attachment to the load cell
5.3.2 Workpiece clamp
The workpiece clamp plate holds the PMMA workpiece in place during
embossing and de-embossing. It should effectively clamp the perimeter of the workpiece
while leaving the middle open for the tool to pass through and contact the workpiece.
The workpiece will be a 1 mm thick PMMA sheet cut larger than the tool to permit
perimeter clamping. The clamp plate will be mounted to the bottom platen, making it
more accessible for replacing the workpiece between runs. The plate will be mounted
74
with four screws fitting tapped holes in the platen. The plate has a recessed area on the
bottom to fit the workpiece, and an oversize hole through which the vacuum chuck can
project. The plate is 0.125 in thick, thus minimizing the height of the vacuum chuck.
The clamp plate will be made of copper for good thermal conductivity.
3J1
Figure 5-15 Workpiece clam plate
5.3.3 Tool chuck (vacuum)
The vacuum chuck must effectively hold a wafer-type tool against the tensile
loads of de-embossing, and also hold the wafer in place during forming. In section 4.5,
the maximum clamping force was found to be about 398 N. The critical feature that
determines the clamping force is the projected area of the grooves in the chuck's surface.
The chuck shown in Figure 5-16 consists of a 0.125 inch plate with a 4 inch circular area
raised 0.125 inches above. The circular area has seventeen concentric grooves that are
0.050 in wide and 0.050 in apart. The projected area of these grooves is 5.12 in2. For a
realistic vacuum of 10 torr (1.33kPa), this gives a theoretical clamping force of 330 N. In
75
practice, leakage and other factors may reduce this value somewhat. The grooves are
connected together by a radial groove that is also 0.050in wide. At the end of this groove
is a 0.050 in diameter hole that penetrates to another hole drilled from the side of the
chuck, providing a flow path to a vacuum fitting as shown in Figure 5-17. The vacuum
chuck will be made of copper for good thermal conductivity.
To effectively hold the clamping vacuum, a compliant gasket between the wafer
tool and the vacuum chuck is necessary. This gasket will be made from a sheet of
Thermagon T-pli 220. This material is a silicone elastomer sheet 0.020 in thick,
reinforced with glass fibers. The elastomer is doped with boron nitride particles to
improve its thermal conductivity. Properties for this material are given in appendix A.5.
Figure 5-16 Vacuum chuck
76
Figure 5-17 Detail of vacuum port
5.3.4 Spacer plate
The vacuum chuck is a total of 0.25 in thick. This added material between the
tool and the top platen will retard heat transfer somewhat, because of both the larger
separation between the tool and to fluid passages and the increase in thermal mass. A
0.25 in spacer plate must be mounted under the workpiece in order to make the top and
bottom platen assemblies symmetrical in terms of heat transfer. This spacer plate will
also allow direct measurement of the workpiece temperature. Small holes can be drilled
through the plate so thermocouples may be put in contact with the PMMA. Small
channels on the underside of the plate would accommodate the thermocouple wires.
77
Figure 5-18 The spacer plate
5.4 Structural and Thermal finite element model of platen
assembly
In order to predict the structural behavior and thermal performance of the platen
assembly, a finite element model was created using ANSYS. A structural analysis was
performed to ensure that the minimal platen design could carry expected embossing loads
without failing, and that the surface of the platen remained uniform under pressure.
Transient thermal analyses of the full platen assembly were performed to assess the
uniformity of heat transfer to the workpiece and tool.
5.4.1 Structural
The structural performance of the platens was evaluated using a finite element
model in ANSYS. The model was a two-dimensional cross section of the platen. A
detail view of the model with the finite element mesh is shown in
78
AN%
ELMNS
APR 25 2005
22:36:51
Figure 5-19 Detail of structural FEA model of platen
This model was had a boundary condition of zero vertical displacement at the
bottom surface, and was subjected to a pressure of 6366 kPa, corresponding to 50 kN
applied over a circular area 100 mm in diameter. A detail view of the resulting von
Mises stress distribution is shown in Figure 5-20. The maximum stress is 15 MPa, well
below pure annealed Copper's yield stress of 33.3 MPa. It is also important that the
platen maintain a uniform surface under load. Figure 5-21 shows the vertical deflection
across the top of the platen. The periodic variation is a result of the fluid passages. The
maximum variation in deflection at the top of the platen is less than 0.035 micron. The
deformation of the platen under load will not have any significant impact on the forming
of the PMMA.
79
NODAL
A I
SOLUTION
APR 25 2005
23:14:26
STEP2=1
SUB
BUB =1
=1
TIME=1
(AVG)
SEQV
DMX =.159E-05
SMN
=435736
SMX =.159E+08
.215E+07
.900E+07
.55E+07
Figure 5-20 von Mises stress in platen (Pa)
80
.141E+08
.107E+08
.729E+07
.386E+07
435736
.124E+08
.159E+08
0.6
2 0.5
E
'0
0.4
L-
CD
= 0.3
0
CD)
%
0.2
CD)
0.1
0.0
0
20
40
60
80
100
Position along platen (mm)
Figure 5-21 Vertical deflection at top of platen
5.4.2 Thermal
It is important that the heat flow to the workpiece and tool during heating and
cooling be as uniform as possible. Differential heating or cooling could result in residual
thermal stresses that could impact part quality. A thermal FEA model of a cross section
of the platen stack was created in ANSYS. The exterior lines were constrained to have
zero heat flux, and the materials were defined as in Figure 5-22. The internal surfaces of
the passages were subject to a convective boundary condition with the convection
coefficient equal to 3500 w/m 2 K and the fluid temperature 25'C. The model was at a
uniform initial temperature of 150'C. The temperature at a node in the center of the
PMMA is shown in Figure 5-23. It can be seen that the settling time is about 60 s,
implying a time constant of about 15 s for the platen assembly.
81
Top Platen
Vacuium Chuck
Gasket
Gasket
Silicon
Tool
Workpiece
\
ppe
Spacer
Bottom Platen
Copper
Figure 5-22 Platen stack model
150
125
100
*1-
75
CL
E
I-
50
25
0
0
10
20
30
40
50
60
Time (s)
Figure 5-23 Temperature in center of PMMA
The temperature distribution across the PMMA workpiece over time is just as
important as the absolute temperature. This distribution is shown as it changes over time
82
in Figure 5-24 Temperature Distribution at bottom of PMMA over time. Some nonuniformity is evident in this figure. Because the fluid tubes at the edges do not have
neighbors, heat transfer at the edges is not as effective as in the center. This phenomenon
is known as the edge effect, and is an inescapable result of the physics of heat transfer.
One extra fluid tube is present on each side of the platens to account for this effect.
83
Time
160
140
-
1 40
T.
-
-
-
-
-
+
-
-
-
-
-
.
-
-
-
-
.-
-
-
-
-
.
-
-
0.5s
0.15s
~-
-
-
-
-
-
-
-
-
-
10
-
-
s
s
-2.0
m
120
4.Os
100
---
0
C-....
-
0
--
-
-.
-
--
-
6.Os
-
---
-
..
-
1s
E
60
-
-
-
-
-
15 s
40
3is
..
i-
--
-..-
-
---
---
-
20
*.
...........................................................................
45 s
60s
>
Width of PMMA workpiece
(100mm)
r)
0.0
20.0
60.0
40.0
80.0
100.0
120.0
Position (mm)
Figure 5-24 Temperature Distribution at bottom of PMMA over time
Initial temperature of 150*C and fluid temperature step change from 150-25*C at time zero
84
0.7
0.6
0.5
0_
0.4
0.3
0.2
0.1
0
0
10
20
30
40
50
60
Time (s)
Figure 5-25 Edge effect over time
The magnitude of the edge effect over time is shown in Figure 5-25. This result
was found by subtracting the average temperature at the middle of the PMMA from the
average temperature at the left edge. It is seen to peak at about 0.6'C after 5 s. The slope
discontinuity at 20 s is a result of changing the time step size from 1 s to 5 s to reduce the
required computation time. The periodicity of the fluid passages also introduces some
slight non-uniformity at the workpiece surface, but the magnitude of this variation is on
the order of 0.02'C, so it is not significant.
There is one additional source of non-uniformity across the platens. As the fluid
flows through the tube, heat is transferred to or from the platen to the fluid. This is, of
course, the function of the fluid. This heat transfer changes the temperature of
the platen,
but it also changes the temperature of the fluid. The situation is most exaggerated when
the difference between the platen temperature and the fluid temperature is large. If the
85
temperature of the platen is 150'C and the temperature of the fluid is 25'C, as would be
the case at the beginning of a fast cooling episode, the temperature of the fluid could
increase by as much as 4.5'C as it passes through the tube. As the platen cools, the
temperature rise in the fluid becomes smaller, as shown in Figure 5-26. The rate of heat
transfer into the fluid depends on the temperature difference by Equation 5-1. This
variation in heat transfer rate along the tube will result in slightly different temperatures
at the surface of the platen in the direction of fluid flow. This temperature variation will
not be larger in magnitude than the variation of the fluid's temperature. This effect can
be mitigated by using less severe differences in temperature between the platen and the
fluid. If a large, fast change in temperature is desired, however, this effect is
unavoidable.
86
30
1
1
I
T=Temperature of platen (C)
2 9 ...----- .-.-..-.---.-..-..-.-.------ ..-.
---..- .----- '-.
1 5-
a=130
..-.----- --. .---- -..-.-.--..--- ..-.-.--.------.. --.
.-.
.-S 2 8 ----..- .-.
4T=110
0
E
CL
....
-.-.-.-.
-. --.-.. -.
..--.
...--.
2 6 ---- --
-----
26
1
0
---.
.--
T=70
-.-.- --.-.-- .-T=50
..-----.------- .---- ..-------.-.------- T= 3 0 .......3
2
Position along tube (in)
4
5
Figure 5-26 Fluid temperature change along a tube
For initial fluid temperature of 25 0C, and platen temperature as shown
5.5 Manifolds
The function of the manifold is to separate the flow of the heat transfer fluid into
the 18 tubes in each platen. The manifolds must be designed to distribute the fluid flow
evenly among the tubes. Various conceptual designs were tested in simulation using the
Flotran CFD module of ANSYS. The best concept proved to be manifolds consisting of
a large bore cylinder perpendicular to the fluid channels. The CFD model for this design
is shown below.
87
Fluid outlet
Fluid inlet
Figure 5-27 Flotran CFD model of manifold design
The CFD model calculated pressure and velocity throughout the manifold and
platen assembly. The outlet of the manifold was constrained to have a pressure of zero,
so the pressure at the inlet is equal to the total pressure drop. This pressure drop is found
to be 121 kPa, or 17.6 psi. The flow velocity data along a line bisecting the fluid tubes
halfway through the platen was extracted, and the average flow velocity in each tube was
calculated. A plot of the flow velocity is shown in Figure 5-28, and the resulting
convection coefficient is shown in Figure 5-29. It can be seen from these plots that the
manifolds do a good job of evenly distributing fluid flow across the platens.
88
I
II
II
9876 -~
0
U--
0
321
1
2
3
4 5
6
7 8 9 10 11 12 13 14 15 16 17 18
Tube #
Figure 5-28 Mean flow velocity in each tube in the CFD model
89
4500
40002 35003300025000
U 2000 0
1500 0
o 10005001
2
3
4
5
6
7
8
9 10 11 12 13 14 15 16 17 18
Tube #
Figure 5-29 Convection coefficient in each tube
Material selection for the manifolds is important in order to minimize their
thermal influence, and to ensure they will not fail mechanically. For obvious reasons of
safely, the manifolds must be guaranteed not to burst at high temperatures. There are
many choices for manifold material. A few polymer materials such as Poly-ether-imide
(PEI) and poly-ether-ether-ketone (PEEK) retain structural properties fairly well at high
temperature; however, the maximum operating temperature for the machine, 200'C, is
nearly out of the acceptable range for these materials. The manifold design shown in
Figure 5-30 was used for a finite element model in Pro/Engineer to determine the stresses
in the manifold under pressure. The model was subject to 100 psi internal pressure,
corresponding to the maximum possible pressure in the fluid supply system (see section
90
6.8). The resulting stress distribution can be seen in Figure 5-3 1. The deformed shape
has been exaggerated.
Figure 5-30 Manifold design
91
Siress vrin Mises (Maximum)
Deformed Or ginol Model
Max Disp +4I.2E-04
&,ale I.-980E+03
4.000e+03
3.500e+03
3. 000e+03
2.500e+03
2. 000e+03
1.500e+03
1.000e+03
P in pal Units:
Inc' Pound Second (IPS)
5.000e42
0.000e 4 00
Figure 5-31 Stress distribution in the manifold
Because of the level of stress present, PEEK and other polymer materials will not
be adequate at the high end of the operating temperature range. Metals retain their
mechanical properties very well in the temperature ranges currently considered. Because
a metal manifold would have stiffness comparable to the copper platen, thermal stresses
must be considered. The area where the platens mate with the manifolds is shown in
cross section in Figure 5-32. The platen and manifold will be joined by brazing.
92
Manifold
Figure 5-32 Thermal stress model
The thermal stress in two dissimilar materials can be found by the following
method. Because the materials are attached together, their final strains must be equal.
The internal forces in the two materials must also be equal. These constraints give
equations Equation 5-10 and Equation 5-11, which may be solved for the internal force as
in Equation 5-12. Parameters with subscript "p" refer to the platen, and subscript "IM"
refer to the manifold. K is the coefficient of thermal expansion, L is the original length,
A is the cross-sectional area, AT is the change in temperature, and , is the elastic strain.
KLAT + e, =KPLAT +c
Equation 5-10
E
EP
'"
Equation 5-11
F=
(K - Km)AT
E A
E, A,
Equation 5-12
For copper a copper platen and a manifold made from 6061 aluminum, the
difference in thermal expansion over a change in temperature from 25 to 150'C will
result in an internal force of 2106 N, corresponding to stresses of 34.8 MPa in the copper
and 104.5 MPa in the aluminum. Both pieces would be yielded. In fact, no other metals
have thermal expansion coefficients close enough to that of copper to reduce thermal
93
stress to acceptable levels. Copper is the only material that is strong enough at high
temperatures and that will not cause significant thermal stress in the platens.
5.6 Insulation
In order to maximize the thermal performance of the platen assembly, it should be
thermally isolated from the rest of the machine. This insulation will also protect the load
cell in the Instron frame's crosshead, which is sensitive to temperature. The load cell is
rated for operation at temperatures from 10 C to 38'C.
Rescor 914 ceramic from Cotronics was chosen for its very low thermal
conductivity, high compressive stress, and machinability. The one-dimensional steadystate heat transfer equation give the heat transfer rate through the ceramic, where q is the
heat transfer rate, k is the thermal conductivity, A is the conducting area, AT is the
temperature difference across the ceramic, and x is the thickness of the ceramic [25].
q=kAAT
x
Equation 5-13
The insulation layer should be sized so that the rate of conduction into the anvil is
no more than the rate of heat loss to the environment. Heat loss from the anvil is
primarily through free convection. Free convection is governed by the Grashof and
Rayleigh numbers. In Equation 5-14 and Equation 5-15, g is the gravitational
acceleration,
P is the thermal expansion coefficient
which is equal to the inverse of
temperature in Kelvin for an ideal gas, Ts is the temperature of the convected surface, T.
is the ambient temperature, L is the characteristic length of the convected surface, v is the
viscosity of air, and a is the thermal diffusivity of air [25].
94
s(TS
-
Gr =2g
T.0 )L3
Equation 5-14
Ra = GrPr=
g/3(T - TJL 3
va
Equation 5-15
The relationship between Nusselt number and the Grashof number is different for
vertical and horizontal surfaces. For a vertical surface, Nusselt number is given by
Equation 5-16. For a hot horizontal plate, Equation 5-17 is valid for the Rayleigh number
in this situation [25].
4 (GrY" 4
Nu = -Pr
I
g
3
4
Equation 5-16
Nu = 0.54Ra1 /4
Equation 5-17
Using these equations with Newton's law of cooling (Equation 5-1) it is found
that the rate of heat loss to the environment is 18.3 W when the anvil is 35'C. With this
value, Equation 5-13 can be solved for the required thickness of the ceramic layer. For
the case where the platens are held at 150'C, the ceramic must be at least 1.97 in thick.
The properties of the ceramic are given in appendix A.6.
5.7 Summary of platen assembly design
The platen assembly is the most critical portion of the HME machine. The
platens must transmit heat to and from the workpiece and tool effectively. The fixturing
components must connect the platens to the load frame and must support and align the
workpiece and tool. The manifolds must distribute fluid flow evenly to the tubes in the
platens for even heating and cooling. Finally, an insulation layer protects the load frame
95
from the heat of the platens, and isolates the thermally cycled components to minimize
the thermal mass. Figure 5-33 shows the main components of the platen assembly.
Figure 5-34 and Figure 5-35 show the full assembly discussed in Chapter 5, including the
mounting hardware and manifolds.
Top Platen
Vacuum
Chuck
Tool
Clamp
MI
-----------------
Workpiece
Spacer go
Bottom Platen
Figure 5-33 Exploded view of platen stack
96
-A" i
Screw block
Screw block
Top canier
9
Insulation
Manifold
~~1
Insulation
I
Manifold
P1lenI
Clamp
1
11 1i
IIIPlatenm
Manifold
Insulation
Insulation
Bottomn
T-Plate
canier
-Spacerl A
I
I
Vacumn cluick
_
I
MEAnfold)
Figure 5-35 Three-dimensional view of the full platen assembly
The thermal mass of each of the platen assembly components subject to thermal
cycling is tabulated below. If it assumed that all components are heated only be
convection inside the platen tubes and that the lumped capacitance assumption applies,
the calculated total thermal mass would give a time constant equal to 21 s, or a settling
time of about 80 s. The average power needed to heat the platen assembly from 25150'C in 80 s is estimated to be 5.4 kW. In practice, the power of the thermal control
system must be greater to heat the circulating fluid in addition to the platens.
98
Component
Manifold (x4)
Platen (x2)
Spacer Plate
Vacuum Chuck
Clamp plate
PMMA
Silicon wafer
Vacuum Gasket
Heat
capacity
J/kg-K
385
385
385
385
385
1450
703
800*
Thermal
mass
J/K
474.3
445.4
330.1
234.0
62.2
20.1
6.4
4.6
Total Thermal Mass:
3446
Volume
Density
Mass
mA3
kg/mA3
kg
1.23E+00
1.16E+00
8.57E-01
6.08E-01
1.62E-01
1.39E-02
9.15E-03
5.72E-03
1.37E-04
1.29E-04
9.57E-05
6.78E-05
1.80E-05
1.17E-05
3.93E-06
4.OOE-06
8960
8960
8960
8960
8960
1190
2329
1430
Figure 5-36 Table of thermal masses (*=estimated property)
99
CHAPTER
6
Design of the temperature control system
6.1 Introduction
The temperature control system serves to circulate the heat transfer fluid through
the system and to control the fluid's temperature. The system will control the
temperature of the heat transfer fluid by modulating the mixing ratio of a hot and cold
fluid stream. This system architecture was shown in Figure 4-2. A more detailed system
diagram is shown in Figure 6-1. The top and bottom platens will have separate control
valves so their temperatures can be set independently. This diagram also includes the
final major component of the fluid system, the expansion tank. This component provides
for the thermal expansion and contraction of the working fluid, provides a reservoir for
excess fluid, provides for venting of gases from the fluid system, and maintains a positive
pressure head at the inlet of the pump.
100
Expansion
Tnk.
Top
Platen
Bottom
Platen
Mixing.
Valves
Heat
Source
L4
......
)
Pum]
Heat
Sink
Figure 6-1 System Architecture
6.2 Selection of oil/water heat exchanger
Plate and frame heat exchangers are compact and powerful, so this type of heat
exchanger will be used to cool the circulating oil. The cooling fluid will be ordinary tap
water. In order to specify a heat exchanger, one must know the flow rate and inlet
temperature for both the oil and the water, and the desired outlet temperature. The
current application is more complex because the flow rate and inlet temperature are not
constant.
As the fluid circulates, it is split between the hot and cold streams, heated or
cooled respectively, and re-combined in the proper ratio to give the desired temperature.
The flow rates of hot and cold fluid needed to produce a desired temperature can be
found by solving the conservation of mass and conservation of energy equations.
101
Qh, Th
Qt, Td
E**
Qc, Tc
Figure 6-2 Mixing hot and cold fluid to produce desired temperature
QhPh +Qp
=Q
QhPhCph (Td
h
tPd
=
- Td
cpc
Equation 6-1
Qh=
QtPdCPC
1
Ph
QC =
PC
PC
c
Td)
-Td )
ph(h -Td
(Qd -QhPh)
Equation 6-2
In the above equations,
Q is flow rate, T is temperature,
p is density, and C, is
specific heat capacity, while the subscript h stands for the hot branch, c for the cold
branch, d for the desired temperature, and t for the total flow rate. By Equation 6-2, the
flow rates in the hot and cold branches can be estimated from the desired temperature,
total design flow rate of 40 gpm, the temperatures of the fluid in each branch, and the
fluid properties. Because the temperature of the fluid does not change very much as it
passes through the platens (see section 5.4.2), the inlet temperature of the heat exchangers
is approximately the same as the desired outlet temperature of the control valves. This
inlet temperature and flow rate information can be used to select a heat exchanger.
102
Figure 6-3 Photo of oil cooler
After consulting with suppliers of compact heat exchangers, the MaxChanger
model MX-22 from Tranter PHE was selected. This model has 4 in wide plates 24 in
long. Figure 6-3 shows a photo of the heat exchanger, and an engineering drawing is
available in appendix B. 11. Performance data for a range of flow rates and temperatures
was provided by the manufacturer. These temperature and flow rate conditions were
estimated using Equation 6-2. The data shows that this heat exchanger is quite capable of
reducing the temperature of the oil significantly in a single pass. For instance, at the 17.9
gpm and 1 00 0 C condition, the effective power of the heat exchanger is 122 kW.
Flow
Pressure
(gpm)
0.12
4.1
11.5
17.9
26.1
31
drop (psi)
0.004
0.286
1.210
2.604
5.589
8.765
Inlet
Outlet
Temp ('C) Temp ('C)
30.4
179
54.6
163
79.7
130
71.2
100
49.8
60
28.9
30
Figure 6-4 Oil cooler performance
Unfortunately, a 121 kW electric heater would not be feasible for a laboratory
instrument. The selection of the heater requires some more detailed knowledge of the
system's thermal performance.
103
6.3 Dynamic thermal model
In order to understand the performance of the proposed temperature control
system, a model was created using MatLab. This model approximates the dynamic heat
transfer behavior in the fluid-based temperature control system. The model includes the
thermal masses of the platens and the electric heating elements within the circulation
heater.
The dynamic model consists of modules corresponding to the platens, the
oil/water heat exchanger, the circulation heater, and the control valves. Each module
uses its inputs at the current time step to calculate its outputs at the next time step. The
inlet temperature and flow rate through the heater, cooler, and platens is used to calculate
their outlet temperatures for the next time step. The outlet temperature of one component
becomes the inlet temperature of the next component in the fluid circuit. The information
flow among the various modules is illustrated in Figure 6-5.
The dynamic model also accounted for the time delay caused by the finite
distance between components. A pure delay of 2 s was inserted between the outlet of the
platens and the inlet of the heat exchangers, corresponding to the time for fluid to travel
through about 2.5 m of 2 in diameter pipe at 40 gpm. A delay of 0.8 s was inserted
between the outlet of the heat exchangers and the inlet of the platens, corresponding to 1
n of such pipe. The delays were implemented by adding the number of time steps
corresponding to the delay to the index of the output variables for the respective
components. For instance, the outlet temperature of the platens was used as the inlet
temperature for the heat exchangers 20 time steps later, rather than just one. These delay
times are based on estimates of the system layout, and are only intended to make the
104
model more representative. The models for each of the components are discussed below.
The MatLab code for the full model and each module is available in appendix C.3.
T
Calculate
r Heat TCet+b
Tr(!)OiVWaW
To(t+)
T80+0
TWO
O
Echne
Tit
Electric
Circulation
Branch Flon
Th()
(Valves)
Platens
Qh(t+I)
-
Tt)
Heater
TWO+1
*r
Tc=Cold branch temp
To=Outlet temp
Tp-Platen temp
Qc-Cold branch flow
Th- Hot branch temp
Ti-Inlet temp
Te-Heating element temp
Qh- Hot branch flow
T d=Desired fluid temp
Figure 6-5 Information flow diagram for the dynamic thermal model.
The current time step is t, the subsequent step is t+1.
6.3.1 Platen model
The dynamic behavior of each thermal mass in the system can be represented by
first-order differential equations. The temperature of the platen assembly is given by
Equation 6-3, where h is the convection coefficient, A is the surface area of the tubes, T,
is the platen temperature, Tm is the fluid temperature, and TM is the thermal mass.
The
convection coefficient is calculated as in section 5.2.1 using Equation 5-3.
d3 _ hA (T - T,
dt
TM
Equation 6-3
105
In the dynamic simulation, this differential equation was solved by Euler
integration. The rate of heat transfer to or from the platens is calculated at each time step
using convection correlations. The heat transfer rate divided by the thermal mass gives
the rate of temperature change. The rate of temperature change is multiplied by the time
step size to give the incremental change in temperature, and this change is added to the
current temperature to give the temperature for the next time step. The time step size was
set to 0.1 s, two orders of magnitude lower than the expected time constant of the platen
assembly. This portion of the MatLab script for modeling the platen assembly is given
below, where q is the heat transfer rate, h is the convection coefficient calculated from
the correlation in Equation 6-6, ConvA is the surface area subject to convection, Tm is
the fluid temperature, Ts is the the platen temperature, TMp is the thermal mass of the
cycled components, dTs is the temperature rate of change, dt is the time step size, and
Tsnew is the platen temperature for the following time step. The full MatLab script is
available in appendix C.3.
%Power gain/loss
from the fluid to the platens
(W)
q=h*ConvA* (Tm-Ts);
TMp=3446;
%Rate of change of Ts (degrees C/sec)
dTs=q/TMp;
%New platen temperature
Tsnew=Ts+dTs*dt;
The temperature of the fluid at the outlet of the platens is found by imposing the
energy balance. The heat transfer rate into the platens must equal the heat transfer rate
out of the fluid, so the fluid outlet temperature is given by Equation 6-4, where T0 and Ti,
are the outlet and inlet fluid temperatures, respectively, q is the rate of heat transfer into
the platens, p is the fluid density, C, is the fluid specific heat, and Q is the fluid volume
flow rate.
106
T"= T. -q
pC~
Equation 6-4
6.3.2 Heater model
Because the electric heater physically resembles a shell-and-tube heat exchanger,
the convective heat transfer from the heating elements to the fluid was estimated using
correlations for the shell-side fluid. The equivalent diameter for fluid calculations for the
shell-side fluid is given by Equation 6-5, where ODT is the diameter of the heating
elements and P 1 is the spacing of the heating elements [33]. The Nusselt number is given
by Equation 6-6.
2
r/4
(p - OD
nODT
Equation 6-5
Nu = 0.36 Re 055 Pr" 3
Equation 6-6
Because the source of heat is electricity, rather than heat loss by another fluid,
equations for the change in fluid temperature developed for shell-and-tube heat
exchangers cannot be used for electric heaters. Instead, the outlet temperature is given by
Equation 6-7, where h is the convection coefficient, A is the surface area of the heating
elements Tt is the temperature of the heating elements, p and C, are the density and
specific heat of the fluid, and
Q is the
fluid flow rate. The temperature of the heating
elements is given by Equation 6-8, where Tt is the heating element temperature, Tm is the
fluid temperature, qE is the rate of Joule heating in the elements, which is equal to the
heater power, and TM is the thermal mass of the heating elements. As with the platen
model, this differential equation is solved using Euler integration.
107
T T +hA(TT-
Tin )
Equation 6-7
dt_ qE- hA(T - T
dt
TM
Equation 6-8
6.3.3 Oil/water heat exchanger model
The manufacturer of the oil/water heat exchanger provided performance data for
various representative combinations of flow rate and inlet temperature. This data was
used to model the heat exchanger in the MatLab script by fitting functions to it. This data
is given in Figure 6-4.
6.3.4 Control valve model
The dynamic response of the control valves and the temperature control feedback
loop was not included in the thermal model. The time response of the final system will
be dominated by the dynamics of the major thermal masses; that is, the platens, heating
elements, and the fluid. The control valves themselves will be able to open and close in
one or two seconds, whereas the time constant for the platen assembly is on the order of
tens of seconds. In the dynamic model, the valves and temperature control feedback loop
were treated as instantaneous and perfect. The resulting variable flows through the hot
and cold branch were determined as in section 6.2 by using Equation 6-2.
For simplicity, there were several phenomena not included in the thermal model.
These include the dynamic response of the control valves, the impact of fluid momentum
on changing flow rates, heat lost to the environment from the fluid system tubing, and
friction heating of the fluid from the pump and other components. From the transient
perspective of the thermal model over the time scale of several minutes, most of these
108
effects can be reasonably assumed to be insignificant. The model as used captures the
most important phenomena and is sufficiently representative to guide the selection of the
heater and to roughly predict the thermal performance of the system.
6.4 Selection of electric circulation heater
Initially, the module of the system simulation that represented the heater had
parameters based on typical circulation heaters in the power range considered. As the
power requirements became more defined, manufacturers of circulation heaters were
consulted to select a specific heater. The final choice is a 30 kW circulation heater from
Vulcan Electric Co. This heater has 18 heating elements within a 44 in long enclosure
with 8 in internal diameter. Each element consists of a U-shaped tube with 0.475 in
diameter, about 33 in long. 3 internal baffles induce swirling motion in the fluid to
enhance heat transfer. Figure 6-6 shows a photo of this heater mounted to a stand, and
drawings are available in appendix B. 12.
Figure 6-6 Photo of circulation heater
109
The dynamic thermal model was configured to simulate them temperature profile
of a typical embossing experiment, comprising an step change in desired temperature
from an initial value to the embossing temperature, and a subsequent step change to a deembossing temperature. Figure 6-7 shows the platen temperature and the desired
temperature results of the dynamic thermal model. The model ran for 550 s before the
data shown on the graph in order to allow the model to reach steady-state conditions after
initialization. The initial and de-embossing temperatures are set to 40'C, and the
embossing temperature is set to 150'C. The time to heat to within 2% of the final
temperature is 88.6 s. The time to cool to within 2% of the final temperature is 75.0 s.
160
I
I
I
-- Platen Temp
--- Desired Temp
140120100
-
80 -60
-..
-..-.
-..
40
20
950
-
600
650
700
750
Time (s)
800
850
900
950
Figure 6-7 Output of dynamic thermal model with the 30kW circulation heater
110
For comparison, the time response with a 25kW heater is show in Figure 6-8. The
time to heat to 2% of final temperature has increased by 54% to 136.3 s. The time
response with a 35kW heater is, shown in Figure 6-9. The time to heat has decreased by
only 19% to 72.0 s. There is a diminishing return for increasing the heater power because
heat transfer to the fluid depends on the temperature difference between the fluid and the
heating elements. The heating elements are constrained to a maximum temperature of
250'C to protect the fluid from thermal degradation, so adding additional power beyond
some point will have no effect at all.
160
1
----
1
1
1
-Platen Temp
--- Desired Temp
---------------
140120100--
E
8060-'--------
40
20--
050
600
650
700
750
Time (s)
800
850
Figure 6-8 Performance with 25kW heater
111
900
950
160
I
I
-Platen Temp
--- Desired Temp
140
-I
120
100
-
I
2) 80
CL
-
I
60
-
I
cI--
E
40
'.----------
20k
I
550
600
650
700
750
Time (s)
800
850
900
950
Figure 6-9 Performance with a 35kW heater
6.5 Predicted dynamic thermal performance
The time response of the system is dependent on the starting and ending points of
the temperature change because the dynamics of the heating element and fluid
temperatures affect the platen temperature. For a large change in temperature, the fluid
must be heated a large amount, so the dynamics of the heater and the working fluid
become important. This is illustrated in Figure 6-10 Temperature dynamics of other
system components and Figure 6-11. These graphs represent the same 40-150'C step
command in temperature as that used for Figure 6-7.
112
When the step change is commanded, the flow rate in the cold branch drops to
zero, while that in the hot branch goes to the maximum, corresponding to the control
valves switching over to the hot branch exclusively. The hot branch temperature
immediately drops because the heater must now work harder, as the flow rate is much
higher. In fact, the heater can not quite keep up with the rate of heat transfer to the fluid,
so the elements are cooled, in effect transferring their stored energy to the fluid. This
phenomenon increases the effective power of the heater considerably, however it is shortlived. As the temperature of the fluid increases, so does that of the heating elements,
until the new equilibrium condition is reached. The control valves now permit a small
amount of flow from the cold branch to produce the desired temperature of 150'C.
When the step change back to 40'C occurs, the flow on the hot branch ceases.
The temperature of the fluid and the heating elements quickly rise to their maximum
values, which will be set by the heater controller in the real system. Meanwhile, the
platens cool until the control valves again mix the flow from the two branches to arrive at
the desired temperature.
113
250
i
i
I-
200
*6
'4,
I
li
-
-
#*
I
DI.~f~y~ T~mv~
a V*
Cold
I
...Cold branch
-- Hot branch
-- Heat elem
I~, 150
0
E
-om
-3
-
-,
100
-
-
mammmmIII
_leIIII
t.=.
11==
...
==11==z
50
..
...
...
...
.
...
...
..
...
II...
....
.....
...
...
...
..
I I.....
r) I
550
I
600
650
700
750
Time (s)
800
850
900
Figure 6-10 Temperature dynamics of other system components
114
950
2.5 x 10
--. Hot branch-mCold branch
2
21
.IRS
*
Ht rac
0.5--
550
600
650
700
800
750
Time (s)
850
900
950
Figure 6-11 Hot and cold branch flow rates
The dynamic model was used to simulate some different step changes in
temperature to illustrate the dependence on starting and ending points. Figure 6-12
shows the system performance for a 90-120*C step. For a smaller change in temperature,
the dynamics of the heater and fluid are less important, so the settling time is shorter.
Figure 6-13 shows the system response to the temperature profile used by Ganesan [17],
beginning at ambient, heating to the embossing temperature of 135'C, then cooling to a
de-embossing temperature of 60*C.
In this graph, there is significant oscillation in the hot and cold branch
temperatures following the step change. This oscillation is an artifact of the model,
caused by the discrete pure delay terms and the model's treatment of fluid temperature.
115
The fluid temperature is calculated at each time step. If there is a significant change in
the flow rate through the heat exchangers, as there is at the step change, there can be a
large, sudden change in the outlet in fluid temperature from one time step to the next.
Because of the pure delays added to the model to represent flow time between
components, this sudden change in temperature takes time to work its way around the
model. When the temperature front again reaches the heat exchangers, it again causes a
sudden change in outlet temperature. In the real system, the fluid will not act as discrete
packets, but will mix as it flows. This mixing will prevent distinct temperature fronts
from existing in the circulating fluid, and will eliminate the potential for oscillation.
25
1
200
I
I
1I
"*
~
~
~
-
~
*~U.**,*-
--
,-
--
--
150
1111111111111111111111-ii e
E
100
-Platen Temp
""Cold branch
--- Hot branch
-- Heat elem
50
r~I
550_
6 0
650
700
750
800
850
Time (s)
Figure 6-12 Temperature dynamics for 90-1201C step
116
900
950
250 - --.-
200--..--
Platen Temp
Cold branch
--- Hot branch
-150 -
..--..--
--- Heat elem
..
100--
5
550
600
650
'-I
700
750
Time (s)
800
I
850
I
900
950
Figure 6-13 Temperature dynamics for 30-135-60*C steps
One of the major goals for the new machine, aside from shortened heating and
cooling times, was the ability to follow arbitrary temperature paths. For instance, it may
be desirable to heat and cool the workpiece using a ramp signal rather than a step change.
Figure 6-14 shows the response of the system to a 0.5 'C/s ramp command from 50'C to
150'C and back. Figure 6-15 shows the tracking error.
117
16C
I
I
I
I
I
I
I
I
I
I
I
I
140 -\
120
\
100 -
E
-
80-
\
60
-
4020
-Platen
Temp
--- Desired Temp
950 400 450 500 550 600 650 700 750 800 850 900 950 1000
Time (s)
Figure 6-14 Ramp response
118
I
I
I
I
I
I
I
I
I
I
I
I
I
I
I
I
I
-
- -
64S 2-0
c.
E
a -2-4-6I
I
I
I
350 400 450 500 550 600 650 700 750 800 850 900 950 1000
Time (s)
Figure 6-15 Ramp tracking error
6.6 Modeling fluid losses
The author's co-workers in the MPCL Grant Shoji and Kunal Thaker created a
MatLab script to determine the fluid pressure losses in the system. It is necessary to
know the system pressure drop in order to select the proper control valves and pump.
Pressure drop in a circular tube or pipe is given by Equation 5-5. Pressure drop in the
heater is given by Equation 6-9, where Nb is the number of baffles in the shell and Ds is
the internal diameter of the shell, De is the equivalent flow diameter from Equation 6-5, p
is the fluid density, V is the fluid flow velocity, g is the gravitational acceleration, and the
friction factor f is given by Equation 6-10 [33]. Data provided by the manufacturer of the
119
oil/water heat exchanger was used for its pressure drop. The Flotran CFD model of the
platen/manifold assembly was run at various temperatures to determine its pressure drop.
Ap =f(Nb+1)D
pv2
De 2g
Equation 6-9
e0.576-0.19 In Re
Equation 6-10
The pressure drop depends on the fluid properties as well as the flow situation, so
it is temperature dependent. This system has the added complication that the flow path of
fluid through the system also depends on temperature, as the control valves mix different
ratios of hot and cold fluid to produce the desired temperature at the platens. The
MatLab script thus had to account for the fluid temperature and the architecture of the
fluid system.
Component
Circulation Heater
Cold Heat Exchanger
Fluid Temperature (C) Into Platens
30
70
110
150
7.9
34.4
84.0
0.0
94.5
54.0
22.7
3.8
Platen/Manifold Assembly
Piping and fittings
100.1
141.9
91.3
101.9
86.4
54.1
85.3
30.3
Total System
336.6
255.2
197.6
203.5
Figure 6-16 Table of component pressure drops (kPa)
6.6.1 Selection of control valves
Valve selection for the system was performed by Grant Shoji. Control valves are
selected based on the intended pressure drop across the valves. This pressure drop is
found from the valve authority, given by Equation 6-11, where VA is the valve authority,
P] is the pressure drop across the valve when fully open, and P 2 is the pressure drop
across the rest of the fluid system [34].
120
VA =
P
+ P2
Equation 6-11
For control valves, it is recommended that the valve authority be between 0.2 and
0.5, with higher values closer to 0.5 considered better [34]. Using the system pressure
drops tabulated in Figure 6-16 and a valve authority of 0.5, the valve pressure drop is
calculated to range from 204 to 337 kPa at 150 and 30*C respectively. The lower of
these two values should be chosen as the design point. Otherwise, the valve authority
would exceed 0.5 at higher temperatures. Commercial valves are specified based on their
flow coefficient, Cv. Once the intended pressure drop across the valve is known, the
needed flow coefficient is found by Equation 6-12, where
Q is the
volume flow rate
through the valve, SG is the specific gravity of the fluid, and AP is the intended pressure
drop [35].
C, =QFA
Equation 6-12
For valves sold in the U.S., the flow coefficient is typically calculated using U.S.
standard units of gallons per minute for flow rate and pounds per square inch for pressure
drop. The desired flow coefficient for the valve is thus 3.083.
After consulting with control valve suppliers, the model 2830 three-way mixing
valve from Warren Controls was selected. The valve is actuated by a Moore model 760
Electro-Pneumatic positioner. The valve and positioner assembly is shown in
121
Figure 6-17 Control valve and positioner
6.6.2 Selection of pump & motor
Pump selection for the system was performed by Kunal Thaker. Two basic types
of pumps are available: centrifugal and positive-displacement. The output of centrifugal
valves varies with the downstream pressure drop. Because the pressure drop in the fluid
system is temperature dependent, the flow output of the pump would be considerably
variable. Positive displacement pumps have constant flow rate, so this type is preferred
for the HME temperature control system.
The pump is selected based on its power and flow rate. The power needed is
calculated by Equation 6-13 where H is the required delivery head, Q is the required flow
rate, p is the fluid density, and g is the gravitational acceleration [33]. Head is calculated
from the Bernoulli equation, rearranged to the form in Equation 6-14, where V is the flow
velocity and z is the change in height for the fluid system.
122
Power =H-Q-p.g
Equation 6-13
H = A--+ V+
2g
pg
z
Equation 6-14
For the current application, only the first term-the static head-is significant.
Taking the combined pressure drops of the system and the control valves at their highest
values (corresponding to the low-temperature condition), the static head is72.6m, giving a
required pumping power of 1.4 kW for fluid density taken at 25'C. The rated power of
the pump must be higher because of friction losses and slip. Slip is caused by fluid
leaking from the high pressure side of the pump back to the inlet side. These factors are
unique to a particular pump design, so the final selection must be done in consultation
with the pump supplier.
The pump chosen for the HME fluid system is the model 3711 gear pump from
Roper Pump Co. The pump is connected through a Roper gear box to a 3-phase 5 hp
motor from WEG Industries. The complete assembly is shown in Figure 6-18. The
electric motor will be controlled using a Hitachi L100 series three-phase inverter.
123
Figure 6-18 Pump, gearbox, and motor
6.7 Sizing of the Expansion Tank
The expansion tank provides a reservoir for the working fluid, and ensures
positive pressure at the pump inlet. It also provides for thermal expansion and
contraction of the fluid during operation of the machine. A rule of thumb for sizing the
expansion tank is that it should be 25% full when the fluid is cold, and 75% full when the
fluid is hot. Using the total volume of the system, the density of the fluid at the cold and
hot temperatures, and conservation of mass, the necessary capacity of the expansion tank
can be estimated. The volumes of the various system components are tabulated below.
Using the system volume without the expansion tank, SV, and the density of the fluid at
high and low temperature, Ph and pc respectively, the expansion tank volume XT is found
by Equation 6-15.
p, (SV + 0.25XT) = m = p, (SV + 0.75XT)
XT = SV
P
-Ph
0.75pa -. 0256p,
Equation 6-15
124
Component
Heater
Cooler
Pump
manifolds
Valves
platens
Piping (Est.)
Expansion Tank
System
. 3
in
1448.2
192.0
57.8
49.5
12.6
2.0
746.9
1039.1
3548.1
Volume
gal
6.27
0.83
0.25
0.21
0.05
0.01
3.23
4.50
15.36
m
2.37E-02
3.15E-03
9.46E-04
8.11E-04
2.06E-04
3.35E-05
1.22E-02
1.70E-02
5.81E-02
Figure 6-19 Table of component volumes
In order to protect the working fluid from oxidizing at high temperatures, a cold
trap will baffle the expansion tank from the atmosphere. The cold trap functions
similarly to the drain trap in a household plumbing system. A diagram of an expansion
tank with a cold trap is shown in Figure 6-20. The trap is physically separated from the
expansion tank so that the fluid in it will remain at a low temperature. The lower tube
allows the fluid level in the expansion tank and trap to equalize, while the upper tube
permits the air in the expansion tank to vent as the fluid expands. The air in contact with
the hot fluid in the expansion tank is isolated from the atmosphere because the upper tube
remains below the fluid level in the cold trap. The air in the expansion tank will become
depleted of oxygen, preventing further oxidation of the hot fluid.
125
Vent
Expansion
Tank
Cold
j Trap
To system
Figure 6-20 Diagram of expansion tank with cold trap
6.8 Safety equipment
Safety is an important priority for any engineered system. A machine capable of
pumping flammable oil at up to 200'C at 40 gpm presents many potential hazards to an
operator. If for some reason the pressure in the system should reach an unsafe level,
measures must be taken to prevent damage. The pump itself has a built-in pressure relief
valve designed to prevent serious damage. The valve will trip if the pressure in the pump
exceeds its rated capacity of 125 psi, and the valve will vent fluid from the high-pressure
discharge side back to the inlet. This valve should not be relied on to protect the rest of
the system from overpressure, as the re-circulation of the fluid through the pump can
cause it to overheat. A pressure relief valve set at 100 psi will be mounted at the outlet of
the pump, the highest pressure area in the system. This valve will vent into the expansion
tank in the case of an overpressure condition. The system components themselves offer a
margin of safety. The pump, heater, heat exchanger, and valves are all rated for at least
125 psi, and the manifolds were designed for 100 psi.
The expansion tank will be fitted with high and low level alarm sensors. The
signals from the sensors will be used to trigger an alarm state in the machine that will
126
shut down the heater and the pump. The high level sensor should prevent overflow from
the expansion tank, and the low level sensor will prevent the pump and heater from
operating when too little fluid is in the system, either because the system has not been
charged or in the event of a catastrophic leak.
6.9 Summary of Temperature control system design
The temperature control system must satisfy two of the most important goals for
the new HME machine. It must reduce the cycle time, and it must be capable of
following user-programmed temperature profiles. The temperature control system will
circulate Paratherm MR heated by a 30kW three phase electric circulation heater and
cooled by a plate and frame heat exchanger using tap water. The temperature of the fluid
will be controlled by diverting fluid through either the heater or the cooler for large
temperature changes, while steady-state temperatures will be controlled by mixing hot
and cold fluid from each branch to produce the desired temperature. Diversion and
mixing will be accomplished using 3-port globe valves with electro-pneumatic
positioners. The fluid will be circulated using a positive displacement pump driven by a
5 hp electric motor. A 5 gal expansion tank will provide a reservoir for excess fluid and
allow for thermal expansion of the fluid during operation. A pressure relief valve at the
pump discharge and high and low level sensors in the expansion tank will ensure safe
operation of the system.
127
CHAPTER
7
Conclusions and future work
7.1 Summary
The market for mass-produced polymer micro-devices is potentially huge [6].
Hot micro-embossing shows great promise for producing micron-scale features in
thermoplastic parts. This process has the advantage of a comparatively small thermal
cycle and high replication accuracy [8]. However, comparatively little work has been
done to characterize HME as a manufacturing process, or to consider issues of process
control [12]. To embark on such a research program will require an embossing machine
that gives the experimenter precise control over all potentially significant process
parameters. These include embossing force, embossing and de-embossing temperatures,
as well as time-domain parameters such as heating and cooling rates, strain rates, and
hold time. Existing machines do not give adequate control over these time-domain
process parameters, and are limited in their total cycle times by long heating and cooling
times. To meet the requirements for a thorough investigation of HME, a new machine is
needed. This thesis has discussed the design process for this new machine.
7.1.1 The final design
The design for the new machine encompasses the platen assembly and the thermal
control system. The platen assembly serves to support and align the workpiece and
embossing tool and to transmit thermal energy to and from them. The design for the new
machine's platens exhibits good thermal uniformity, even during transient conditions
128
such as heating and cooling. This is accomplished by having many small, closely spaced
fluid passages. Fixing the workpiece in the machine will be improved through the use of
a vacuum chuck. Lateral alignment of the bottom and top platens is adjustable, and the
mechanism for adjustment has been improved over the existing system.
The temperature of the circulating fluid will be controlled by mixing hot and cold
streams using electro-pneumatic three-port globe valves. The hot stream is heated by a
30kW electric circulation heater, and the cold stream is cooled by tap water in a plateand-frame heat exchanger. The fluid is circulated at 40 gpm by a 5hp positive
displacement pump.
7.1.2 Predicted system performance
Using a dynamic numerical simulation, the thermal performance of the
temperature control system can be estimated. The heating time for a typical embossing
cycle consisting of a step change from 40'C to 150'C is less than 90 s, and the cooling
time is less than 80 s. The temperature control system is capable of following userprogrammed temperature profiles such as heating and cooling ramps as well.
7.2 Conclusion
The design discussed above is not an optimal design, nor was it intended to be.
Without a realistic metric with which to compare the "cost" of increasing system
complexity, whether because of higher pressures or power consumption, with the
"benefit" of improved performance, subjective judgment took the place of quantitative
deduction for many important design decisions. The ultimate goal for this project was
not to hit a specific target, say heating time less than two minutes, but to design a new
machine that would expand the envelope for experiments on hot micro-embossing. The
129
raw performance of the machine in terms of maximum forces and temperatures and fast
cycle time is important, but is secondary to the capability to precisely and repeatable
control the process parameters. The new design certainly gives experimenters access to
faster heating and cooling times to probe the ultimate limits of the HME process, but it
also enables them to specify a particular heating or cooling regime. The new machine
also incorporates improvements in fixturing and alignment of the workpiece and tool.
The faster cycle times and integrated control of the new machine will also make
experiments more convenient.
There are many unanswered questions regarding the hot micro-embossing
process, such as what process parameters are significant, what are their optimal values,
how do these values change depending on workpiece material, tooling material, or
desired geometry, and what is the relationship between disturbances in parameters and
final part quality? When complete, the new HME machine will give its users the
capability to answer these and many other questions. The path to mass production of
polymer micro-devices will be prepared by establishing a firm scientific understanding of
hot micro-embossing.
7.3 Future work
The new HME machine is much more complex than the one it replaces. Design,
analysis, simulation, machining, and the vicissitudes of equipment procurement have
consumed uncountable man-hours. At the time of writing, much of the fabrication work
for the new machine remains to be done. Some components of the platen assembly need
to have machining work finished. The major components of the temperature control
system have arrived, but the specific layout of fittings and pipes is ongoing. Detailed
130
design of the expansion tank and cold baffle also remains to be done. An important and
significant addition will be the control program and electronics that will actually manage
the integrated functioning of the system.
Once it is fully assembled and operating, the machine itself should become the
subject of experiments. The projected thermal performance as discussed in section 6.3
should be compared with the performance of the real machine. The projected thermal
uniformity experienced by the workpiece should also be verified. The capability of the
machine and its attendant control hardware and software to accurately track userprogrammed temperature, force, and displacement profiles should be evaluated.
The platen assembly components have been machined with great care; however,
the typical machine shop precision of +0.001 inch equals 25.4 microns. A 25 micron
parallelism or flatness error in a machine that is intended to form micron-scale features
may be significant, so an active micro-alignment system would be a beneficial addition to
the machine.
131
Appendix
132
A
I
A.1
Material properties
Properties of Paratherm MR
From Paratherm product web page:
http://www.paratherm.com/Paratherm-MR/mr-thermal-oil.asp
Temperature
Density Viscosity
Vapor
Thermal
Specific
Heat
Conductivity Pressure
kPa
W/(m-K)
J/Kg-K
2077.5
0.1477
2091.3
0.1473
0.1468
2104.7
0.1464
2118.1
0.1459
2131.9
0.1455
2145.3
0.1451
2159.1
0.1446
2172.5
OC
-20
-15
-10
-5
0
5
10
15
kg/m^3
20
807.6
5.51
2185.9
0.1442
25
30
35
40
45
50
55
60
65
70
75
80
85
90
95
100
105
110
115
120
125
130
135
803.6
799.6
795.5
791.5
787.5
783.5
779.4
775.4
771.4
767.4
763.4
759.3
755.3
751.3
747.3
743.2
739.2
735.2
731.2
727.2
723.1
719.1
715.1
4.79
4.19
3.70
3.29
2.94
2.65
2.39
2.18
1.99
1.82
1.68
1.55
1.44
1.34
1.25
1.17
1.10
1.04
0.977
0.923
0.874
0.830
0.789
2199.7
2213.1
2227.0
2240.4
2253.8
2267.6
2281.0
2294.8
2308.2
2321.6
2335.4
2348.8
2362.6
2376.0
2389.4
2403.2
2416.6
2430.4
2443.8
2457.2
2471.0
2484.4
2498.3
0.1438
0.1433
0.1429
0.1425
0.1420
0.1416
0.1412
0.1407
0.1403
0.1398
0.1394
0.1390
0.1385
0.1381
0.1377
0.1372
0.1368
0.1364
0.1359
0.1355
0.1351
0.1346
0.1342
839.8
835.8
831.7
827.7
823.7
819.7
815.6
811.6
mPa-s
26.7
20.7
16.4
13.2
10.8
8.96
7.54
6.41
133
0.01
0.01
0.01
0.02
0.02
0.03
0.04
0.04
0.07
0.10
0.12
0.15
0.18
0.26
0.34
140
711.1
0.752
145
150
155
160
165
170
175
180
185
190
195
200
205
210
215
220
225
230
235
240
245
250
255
260
265
270
275
280
285
290
295
300
305
310
316
707.0
703.0
699.0
695.0
691.0
686.9
682.9
678.9
674.9
670.8
666.8
662.8
658.8
654.8
650.7
646.7
642.7
638.7
634.7
630.6
626.6
622.6
618.6
614.5
610.5
606.5
602.5
598.5
594.4
590.4
586.3
582.3
578.2
574.2
570.1
0.717
0.685
0.656
0.629
0.603
0.580
0.558
0.537
0.518
0.500
0.484
0.468
0.453
0.439
0.426
0.413
0.402
0.391
0.380
0.370
0.361
0.351
0.343
0.335
0.327
0.319
0.312
0.305
0.299
0.295
0.291
0.287
0.283
0.279
0.276
134
2511.7
2525.1
2538.9
2552.3
2566.1
2579.5
2592.9
2606.7
2620.1
2633.9
2647.3
2660.7
2674.5
2687.9
2701.7
2715.1
2728.5
2742.4
2755.8
2769.6
2783.0
2796.4
2810.2
2823.6
2837.0
2850.8
2864.2
2878.0
2891.4
2904.8
2919.0
2932.9
2946.7
2960.5
2974.3
2988.1
0.1337
0.1333
0.1329
0.1324
0.1320
0.1316
0.1311
0.1307
0.1303
0.1298
0.1294
0.1289
0.1285
0.1281
0.1276
0.1272
0.1268
0.1263
0.1259
0.1255
0.1250
0.1246
0.1242
0.1237
0.1233
0.1228
0.1224
0.1220
0.1215
0.1211
0.1207
0.1202
0.1198
0.1194
0.1189
0.1185
0.42
0.51
0.59
0.71
0.83
0.96
1.08
1.20
1.60
1.99
2.39
2.78
3.18
4.06
4.95
5.84
6.72
7.61
9.45
11.30
13.10
15.00
16.80
20.40
24.00
27.60
31.30
34.90
41.70
48.50
64.20
80.00
95.80
111.60
127.40
143.20
A.2
Properties of PMMA
From MatWeb page on PMMA:
http://www.matweb.com/search/SpecificMaterial.asp?bassnum=O1303.
The site included the following disclaimer: "The property data has been taken
from proprietary materials in the MatWeb database. Each property value reported is the
average of appropriate MatWeb entries and the comments report the maximum,
minimum, and number of data points used to calculate the value. The values are not
necessarily typical of any specific grade, especially less common values and those that
can be most affected by additives or processing methods."
Metric
English
Comments
Density
1.19 - 1.2 g/cc
0.043 0.0434 lb/in3
Water Absorption
0.13
Average = 1.19 g/c c;
Grade Count = 4
Average = 0.22%;
Grade Count =4
Physical Properties
Water Absorption at
Saturation
-
0.35 %
0.13 - 0.35 %
1.1 %
1.1 %
Mechanical Properties
Hardness, Barcol
49
49
Hardness, Rockwell M
90 - 94
90 - 94
60 - 83 MPa
8700 - 12000 psi
60 MPa
8700 psi
Tensile Strength,
Ultimate
Tensile Strength, Yield
Elongation at Break
4.2
5.5 %
4.2 - 5.5 %
Tensile Modulus
2.8 - 3 GPa
406 - 435 ksi
Flexural Modulus
3 - 3.3 GPa
435 - 479 ksi
-
135
Grade Count
1
Grade Count = 1
Average = 92; Grade
Count = 2
Average = 73.2 MPa;
Grade Count =4
Grade Count = 1
Average = 4.8%;
Grade Count = 4
Average = 2.9 GPa;
Grade Count = 2
Average = 3.2 GPa;
Grade Count = 2
Flexural Yield Strength
100 - 114 MPa
Compressive Yield
Strength
100 - 124 MPa
Izod Impact, Notched
0.22 - 0.25 J/cm
14500 16500 psi
1450018000 psi
0.412 - 0.468 ftlb/in
Average = 110 MPa;
Grade Count = 2
Average = 110 MPa;
Grade Count=2
Average = 0.23 J/cm;
Grade Count = 2
Electrical Properties
Electrical Resistivity
Surface Resistance
le+015 -
1.6e+016 ohm-cm
1.9e+015 ohm
le+015 1.6e+016 ohmcm
1.9e+015 ohm
Dielectric Constant
2.7-4
2.7-4
Dielectric Constant,
Low Frequency
3.5-4
3.5-4
Average = 9E+ 15
ohm-cm; Grade
Count = 2
Grade Count = 1
Average = 3.3; Grade
Count = 2
Average = 3.8; Grade
Count = 2
Dielectric Strength
17 kV/mm
432 kV/in
Grade Count = 2
Dissipation Factor
0.02 - 0.055
0.02 - 0.055
Average = 0.038;
Grade Count = 2
Dissipation Factor,
Low Frequency
0.055 - 0.06
0.055
-
0.06
Average = 0.057;
Grade Count = 2
Thermal Properties
CTE, linear 20'C
Heat Capacity
Thermal Conductivity
Maximum Service
Temperature, Air
Deflection Temperature
at 1.8 MPa (264 psi)
Vicat Softening Point
Minimum Service
Temperature, Air
Glass Temperature
Average = 98.3
61 - 130 pm/m-*C
33.9 72.2 3in/in-0 F
1.5 J/g- 0C
0.359 BTU/lb- 0 F
Grade Count = 3
0.19 - 0.25 W/mK
1.32 - 1.74 BTU-
Average = 0.2 W/m-
in/hr-ft _OF
K; Grade Count = 4
65 - 112 0C
149 - 234 F
Average = 94.5'C;
99 - 112 0C
210 - 234 F
110 OC
230 OF
Grade Count = 4
Average = 100 C;
Grade Count=4
Grade Count = 1
-40 0C
-40 OF
Grade Count = 1
100 OC
212 OF
Grade Count = 1
Optical Properties
136
2
ptm/m-0 C; Grade
Count=3
Refractive Index
Haze
Transmission, Visible
Processing Properties
Processing
Temperature
1.49
1.49
0.6 - 1 %
0.6-1%
92 %
92 %
180 C
356 OF
137
Grade Count = 1
Average = 0.77%;
Grade Count = 3
Grade Count = 3
Grade Count = 1
A.3
Properties of Copper
From MatWeb page on Oxygen-free copper, UNS C 10100.
http://www.matweb.com/search/SpecificMaterial.asp?bassnum=MC101A
Metric
English
8.89 - 8.94 g/cc,
0.321 0.323 lb/in 3
75 - 90
90-105
75 - 90
90-105
221 - 455 MPa
32100 - 66000 psi
69 - 365 MPa
10000 - 52900 psi
55 %
115 GPa
0.31
55 %
16700 ksi
0.31
Machinability
20 %
20 %
Shear Modulus
44 GPa
6380 ksi
1.7le-006 ohm-cm
1.71e-006 ohm-cm
CTE, linear 20'C
17 ptm/m- 0 C
9.44 ptin/in- F
CTE, linear 100'C
17.3 ptm/m-0 C
9.61 gin/in-*F
CTE, linear 250'C
17.7 gm/m-*C
9.83 pin/in-0 F
Heat Capacity
0.385 J/g-0 C
0.092 BTU/lb- 0 F
2660 - 2710 BTU-
Comments
Physical Properties
Density
Mechanical Properties
Hardness, Vickers
Hardness, Vickers
Tensile Strength,
Ultimate
Tensile Strength,
Yield
Elongation at Break
Modulus of Elasticity
Poisson's Ratio
Electrical Properties
Electrical Resistivity
2 hard
full hard
Varies with heat
treatment.
Varies widely with
heat treatment.
in 101.6 mm (4 in.)
UNS C36000 (freecutting brass)= 100%
at 200 C (68 0 F)
Thermal Properties
Thermal Conductivity
Melting Point
383 - 391 W/m-K
in/hrft20 F
1980 OF
1083 *C
138
from 20-100'C (68212 0F)
from 20-200 0 C (68390 0 F)
from 20-300 0 C (68570 0 F)
at 20 0 C (68 0 F)
at 20 0 C (68 0 F)
A.4
Properties of Silicon
From MatWeb page on elemental silicon:
http://www.matweb.com/search/SpecificMaterial.asp?bassnum=MESiO0
Physical Properties
Density
a Lattice Constant
Volume compressibility, 10A10 m 2 /N
Metric
English
2.329 g/cc
5.43072 A
0.306
0.0841 lb/in3
5.43072 A
0.306
11270
11270
112.4 GPa
120 MPa
98.74 GPa
0.28
43.9 GPa
16300 ksi
17400 psi
14300 ksi
0.28
6370 ksi
0.01 ohm-cm
-3.90E-06
0.01 ohm-cm
-3.90E-06
6.7 - 7.1 K
6.7 - 7.1 K
11.8
1.107 eV
1900
500
11.8
1.107 eV
1900
500
1800 J/g
2.49 pm/m-*C
3.61 pm/m- 0 C
4.15 pm/m-0 C
4.44 ptm/m- 0 C
0.702 J/g-0 C
124 W/m-K
1412 C
324 kJ/mol
774 BTU/lb
1.38 pin/in- 0F
2.01 pin/in-0 F
2.31 ptin/in- F
2.47 pin/in-0 F
0.168 BTU/lb-*F
861 BTU-in/hr-ft 2-OF
2570 OF
324 kJ/mol
Comments
Mechanical Properties
Knoop Microhardness
Modulus of Elasticity
Compressive Yield Strength
Bulk Modulus
Poisson's Ratio
Shear Modulus
Electrical Properties
Electrical Resistivity
Magnetic Susceptibility
Critical Superconducting Temperature
Dielectric Constant
Band Gap
Electron Mobility, cm 2 /V-s
Hole Mobility, cm 2 /V-s
Thermal Properties
Heat of Fusion
CTE, linear 20'C
CTE, linear 250'C
CTE, linear 500'C
CTE, linear 1000'C
Heat Capacity
Thermal Conductivity
Melting Point
Heat of Formation
139
N/mm
2
microhardness
Calculated
Atomic (cgs)
6.7-7.1 K, 12.013.0 GPa pressure
at 250 C
at 227 0C
at 527 0C
at 1027 0 C
Debye Temperature
Optical Properties
Refractive Index
Reflection Coefficient, Visible (0-1)
Descriptive Properties
Crystal Structure
372 0 C
702 OF
3.49
0.3-0.7
3.49
0.3-0.7
Cubic
140
at 589 nm
varies irregularly
with wavelength.
Diamond Structure - Space Group Fd3m
A.5
Properties of Thermagon T-Pli 220
From Thermagon product web page:
http://www.thermagon.com/pdf/t-pli200.pdf
Boron Nitride filled, Silicone
CONSTRUCTION/COMPOSITION
CONSRUCTON/CMPOSTIONElastomer, Fiberglass optional
COLOR
THICKNESS
THICKNESS TOLERANCE
DENSITY
HARDNESS
TENSILE STRENGTH
ELONGATION %
OUTGASSING TML (POST CURED)
OUTGASSING CVCM (POST CURED)
UL FLAMMABILITY RATING
SHELF LIFE
TEMPERATURE RANGE
Blue
0.02in (0.508mm)
+ 0.002in (0.05mm)
1.43 g/cc
70 Shore 00
35 psi
5
0.07%
0.02%
94 HB
Indefinite
-45 to 200 0C
THERMAL CONDUCTIVITY
THERMAL IMPEDANCE @20psi
6 W/mK
0.21 oC-in 2/W
THERMAL IMPEDANCE @138 Kpa
COEFFICIENT OF THERMAL EXPANSION
BREAKDOWN VOLTAGE
1.35 oC-cm 2 /W
123 ppm/C
4000 Volts AC
5x1013 ohm-cm
3.26
<0.001
VOLUME RESISTIVITY
DIELECTRIC CONSTANT @ 1MHz
DISSIPATION FACTOR @ 1MHz
141
T-pli 220-AO
40
-
-
36
32 -28
-
2420 C'
o16
12
0
8
4
0
10
20
30
60
50
40
Compression (psi)
70
80
90
100
T-pli 220-AO
40
--
3632-0
~28
2424-
2016-
1280
0
100
200
400
300
Compression (KPa)
142
500
600
700
A.6
Properties of Rescor 914 Glass Ceramic
From Cotronics product web page:
http://www.cotronics.com/vo/cotr/pdf/914.pdf
Use Temperature OF (Max.) 1000
Compressive Strength (psi) 40,000
Flexural Strength (psi) 26,000
Thermal Expansion (x 10-6 / OF) 5.2
Thermal Conductivity (BTU-in / Hr "F Ft2) 2.8
Density (gm/cc) 2.6
Dielectric Strength (volt/ mil) 480
Resistance (ohm/cm) 1014
Loss Factor (@ 1 Mhz) 0.01
Dielectric Constant (@ 1 Mhz) 7.5
143
B
Component drawings
Unless otherwise noted, all dimensions in this section are inches.
B.1
Platen
6. 125
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This part was designed in mm in order to match the top anvil in the Instron, which
was dimensioned in mm.
0
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Dimensions in mm. This part was designed by Ganesan [17].
5
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Face lop & botlome fHui & parol lel, Final ihickness no less ihon 34mm
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S crew block
Top cartier
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MM
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ELECTRIC
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NOTES :
1. RAFFLES JSF BE PLACED H A NAMER N MCH
THERMOCOUPLE IS NOT CGSTROCTED.
2. SRMAP ON RMCE BELOW IN COHDUIT HUB: WLCM.
WATIS, bVLI5, PHAME AND DATE CDE.
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PAMIT COVER AND HOUING WTH RED PAINT.
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C.1
MatLab code
Properties of Paratherm MR
This function uses functions that were fit to the property data for Paratherm MR.
The function returns the properties of Paratherm at any arbitrary temperature.
function [rho,mu, cp, k]=props (T)
%Sets porperties of Paratherm MR at given Temp in deg C for T:
0<=T<=250
rho=-.80441*T+823.69;
mu=.001*(1.0945E-12*T^6 - 9.6662E-10*T^5 + 3.4202E-07*T^4 - 6.2338E05*TA3 + 6.3028E-03*TA2 - 3.5758E-01*T + 1.0649E+01);
cp=2.713*T+2131.8;
k=-8.714e-5*T+.14594;
C.2
Parametric model of internal convection for platens
This function returns vectors containing important performance data for the
platens based on the tube diameter and the flow velocity, given the temperature of the
fluid and the initial temperature of the platens.
function
[DelT,Re,h,Pdrop,tau,DD,VV]=convection
th(Tm,Ts)
%Internal convection model
%Units
are SI
%D=tube diameter L=tube length V=flow velocity
%Tm=mean temp of fluid Ts=tube surface temp
%Change diameter by increments
for diam=1:20
D=diam*.001/2+.002;
DD (diam)=D;
%Calculate platen characteristics based on D
nchan=round(4*.0254/D)+2; %number of tubes per platen
L=nchan*D*2+D; %Length of platen
VolP=3*D*L*L-nchan*pi/4*D^2*L; %Volume of platen
ConvA=nchan*pi*D*L; %Convective surface area
%Change flow velocity by increments
for vel=1:30
V=vel/2+5;
156
VV(vel)=V*pi/4*D^2;
%Call convection model
[DelT(diam,vel),
Re(diam,vel),
h(diam,vel),
Pdrop(diam,vel)]
cvf(D,V,L,Tm,Ts);
%Calculate time constant
tau(diam,vel)=8960*VolP*385/(h(diam,vel)*ConvA);
end %vel
end %diam
end %func convectionth
function
[DelT,Re,h,Pdrop]
%Convection correlation
%Units
are
= cvf(D,V,L,Tm,Ts);
for Turbulent internal
flow
SI
%Find fluid properties at Mean fluid inlet temp
[rho,mu, cp, k]=props (Tm);
P=pi*D; %Perimeter of tube
Vdot=V*pi*(D/2)^2; %Volume flow rate
mdot=V*pi*(D/2)^2*rho; %Mass flow rate
%Dimensionless quantities
Re=V*D*rho/mu;
Pr=cp*mu/k;
%Select proper correlation
if Re<2300
f=64/Re;
Nu=3. 657;
for laminar or turbulent
flow
else
f=(1.58*log(Re)-3.28)^-2;
Nu=(f/2)*(Re-1000)*Pr/(1+12.7*(f/2)^.5*(Pr^(2/3)-1));
end %if
h=Nu*k/D; %Convection coefficient
%Change in fluid temp along tube
DelT=(Tm-Ts) *exp(-h* (P*L) / (cp*mdot) )+Ts-Tm;
Pdrop=f*L/D*V^2*rho/2; %Pressure drop along tube
end %func cvf
C.3
C.3.1
Dynamic thermal model
Main program
The main program sets up the command temperature profile and calls the
functions representing the components.
%Dynamic simulation for platens and heat exchangers.
%Td=command temp
%Tin=input temp to HXs Tc=output temp of CHX Th=output temp of HHX
Tt=temp of HHX coil
%Tvo=output temp of control valve Tpi=input temp to platen Tpo=output
temp of platen
157
=
%Tmax=max temp of coil
%Qp=flow thru platen Qc=flow thru cold branch Qh=flow thru hot branch
disp('go');
%Set up model
flag=4;
watt=30E3;
Tmax=250;
%Sets heater
power
and max temp
Qp=.0024; Pin=200000; %Sets design flow rate Pin is needed as an input
for other functions, but pressure is not included in this
simulation.
dt=.1; %Time step size in seconds
delayv=8; delayp=20; %delayv=delay from valve to platen, delayp=from
platen to HEXs
%set
up command temp profile
len=10000;
%len steps long
initial=80; %init temp
levell=100;
tl=6000; %first
command temp (levell)
and time
level2=80; t2=tl+2000; %second command temp (level2) and time
%generate command temp profile
for i=1:(tl-l)
Td(i)=initial;
end
for
i=tl:(t2-1)
Td(i)=levell;
end
for i=t2:len
Td(i) =level2;
end
comgen=Td;
%Generate time variable for plotting
for
i=l:len
Time (i)=i/dt;
end
%Initial conditions
Th(1)=initial;
Tc (1) =initial-
1;
Ts(l)=initial;
Tt(l)=initial;
Tin(1:l+delayp)=initial;
Tpi(1:l+delayv)=initial;
Qc(l)=Qp/2; Qh(1)=Qp/2;
% Main Program loop
for t=l:len
%Find output of cold HX
[Pout,Qout,Tc(t+1)]
= coldhex new(Pin,
Qc(t),
Tin(t));
%Find output of hot HX
[Pout,Th(t+1), Tt(t+1), fpow(t),
cpow(t)]=shelltubedyn2(Qh(t),Tin(t),Tt(t),watt,Tmax,dt);
tpow(t)=fpow(t)+cpow(t);
%Find outlet temp and new Current temp for platen
[Pout,Qout,
Tpo(t+l),
Ts(t+1)]
= platen dyn(Pin,
Ts(t),dt);
%Pipe delay between platen and HX
Tin(t+1+delayp)=Tpo(t+1);
%Find new flows
[Qh(t+1),
Qc(t+1),
Tvo(t+1)]=tempratios2(Td(t),Qp,Th(t),Tc(t));
%Pipe delay between valve and platen
Tpi(t+1+delayv)=Tvo(t+1);
end
158
Qp,
Tpi(t),
C.3.2
Cold heat exchanger module
This function is based on data supplied by the heat exchanger manufacturer.
function
global t
[Pout,Qout,Tout] = coldhexnew(Pin,
%Pdrop=pressure drop,
Tout=outlet
Qin,
Tin)
temp
%Pdrop and Tout from functions fitted to data provided by maxchanger
Pdrop=1.6481ElO*Qin^2-2.2281E6*Qin;
Tout=-6.3296E-5*Tin^3+1.1668E-2*Tin^2-7.7929E-3*Tin+20.744;
[rhoi,mui,cpi,ki]=props(Tin);
[rhoo,muo,cpo,ko]=props(Tout);
Qout=Qin*rhoi/rhoo;
Pout=Pin-Pdrop;
C.3.3
Electric heater module
This function is based on the convection correlation for a shell and tube heat
exchanger.
function
[Pdrop,
Tout,
Ttnew,
fpow,
cpow]
= shelltubedyn(Qin,
W, Tmax,dt)
%Qin=.001; Tin=100; W=20E3;
%Simulates shell & tube HX where tubes are electric elements.
%Patterned from Vulcan 30 kW heater
Tm=Tin;
[rhom,mum,cpm,km]=props(Tm);
%HEX characteristics
L=(32+11/16)*.0254; %Length of shell
Nb=3; %Number of baffles
B=L/Nb; N=36; %Baffle length (1=none), Number of tubes
ODt=.475*.0254; %Diameter of tubes
C=.145*.0254; %Separation of tubes
Ds=8*.0254; %Diameter of shell
Pt=ODt+C; %Tube pitch
De=3.46*Pt^2/(pi*ODt)-ODt;
As=Ds*C*B/Pt; %Characteristic
%Fluid correlations
Vs=Qin/As; %Fluid velocity
Re=Vs*De*rhom/mum;
Pr=cpm*mum/km;
Nu=.36*Re^.55*Pr^(1/3);
%Effective Diameter
flow area of shell
159
Tin,
Tt,
h=Nu*km/De;
Ao=L*pi*ODt*N; %Convective area of tubes
%Tube thermal dynamics;
mt=.59*L*N; %5.44 kg total mass of heat elements
cpt=481; %Cp of carbon steel
dTt=(W-h*Ao* (Tt-Tm))/(mt*cpt); %Tempurature rate
Ttnew=Tt+dTt*dt; %New temp
%Overlimit setting
if Ttnew>Tmax
Ttnew=Tmax;
end
Tout=Tin+ (h*Ao* (Ttnew-Tm) )/(Qin*rhom*cpm);
%Thermostat setting
of change
if Tout>180
Tout=180;
end
ICalculate
power going into fluid fpow and power going into heating
elements cpow
fpow=(Tout-Tin)*Qin*rhom*cpm;
cpow=W-h*Ao*(Tt-Tm);
%Pressure drop
f=exp(.576-.19*log(Re));
Pdrop=f*(Nb+1)*Ds/De*rhom*Vs^2/2;
shelltube
C.3.4
dyn= [Pdrop, Tout, Ttnew, fpow, cpow];
Platens
Based on parametric model of platens
function
[Pout,Qout,Tout,Tsnew]
=
platen
dyn(Pin,
Q,
Tm,
[rho,mu, cp, k]=props (Tm);
D=.003175;
L=. 132588;
%Flow
V=Q/(36*pi/4*D^2);
mdot=Q*rho;
%Dimensionless quantities
Re=V*D*rho/mu;
Pr=cp*mu/k;
if
Re<2300
f=64/Re;
Nu=3.66;
disp('IN NONTURB');
else
f=(.790*log(Re)-1.64)^-2;
Nu=((f/8)*(Re-1000)*Pr)/(1+12.7*(f/8)^.5*(Pr^(2/3)-1));
end
%Heat transfer
h=Nu*k/D;
160
Ts,dt)
Tout= (Tm-Ts) *exp (-h* (pi*D*L) / (cp*mdot) ) +Ts;
[rhoo,muo,cpo,ko]=props(Tout);
Qout= (Q*rho)/rhoo;
%Find pressure drop
Pdrop=f*L/D*V^2*rho/2;
Pout=Pin-Pdrop;
%Platen convective area
ConvA=L*pi*D*36;
'6Power gain/loss from the fluid to the platens
(W)
Q=h*ConvA*(Tm-Ts);
TMp=3446;
%Rate of change of Ts
(degrees C/sec)
dTs=Q/TMp;
Tsnew=Ts+dTs*dt;
platendyn=[Pout,Qout,Tout,Tsnew];
C.3.5
Calculate branch flows
function [Qh,Qc,To]=tempratios(Tp,Qp,Th,Tc)
%Calculates flowrates of hot and cold sides to produce
%desired temperature and flow
%Qp=flowrate through platen Tp=temp of fluid to platen
%Get properties
if Th==O
Th=180;
end
if Tc==O
Tc=25;
end
[rhoc,muc,cpc,kc]=props(Tc);
[rhoh,muh,cph,kh]=props(Th);
[rhop,mup,cpp,kp]=props(Tp);
%Calculate flowrates
Qh= (Qp*rhop*cpc* (Tp-Tc)
)/(rhoh*cph*
(Th-Tp) +rhoh*cpc* (Tp-Tc));
Qc=(Qp*rhop-Qh*rhoh)/rhoc;
low=lE-12;
if Qh<low
Qh=low;
elseif Qh>Qp
Qh=Qp;
end
if Qc<low
Qc=low;
elseif Qc>Qp
Qc=Qp;
end
Qtm=(Qh*rhoh+Qc*rhoc)/rhop;
161
To= ((Qh*rhoh*cph*Th) + (Qc*rhoc*cpc*Tc)
if To>Th
To=Th;
elseif To<Tc
To=Tc;
end
tempratios=[Qh,Qc,To];
162
)/(Qtm*rhop*cpp);
References
1 A. Manz, N. Graber and H. M. Widmer, "Miniaturized total chemical analysis systems:
A novel concept for chemical sensing," Sensors andActuators B: Chemical,
Volume 1, Issues 1-6, January 1990, p. 244-248.
2 A. Manz, J. C. Fettinger, E. Verpoorte, H. Luedi, H. M. Widmer, D. J. Harrison,
"Micromachining of monocrystalline silicon and glass for chemical analysis
systems. A look into next century's technology or just a fashionable craze?"
Trends in Analytical Chemistry, v 10, n 5, May, 1991, p 144
3 P. Gravesen, J. Branebj erg, 0. S. Jensen, "Microfluidics - a review," Journalof
Micromechanicsand Microengineering,v 3, n 4, Dec, 1993, p 168-182
4 M. Whitfield, "From technology to market," Chemistry World. Vol. 1, No. 12,
December, 2004
5 A. de Mello, "Plastic fantastic?" Lab on a Chip. Issue 2, 2002, p 3 1N-36N
6 H. Becker, L. E. Locascio, "Review: Polymer microfluidic devices." Talanta. Vol
56, 2002, p 267-287.
7 A. J. Ricco, T. D. Boone, N. H. Fan, I. Gibbons, T. Matray, S. Singh, H. Tan, S. J.
Williams. "Application of disposable plastic microfluidic device arrays with
customized chemistries to multiplexed biochemical assays," Biochemical Society
Transactions. Vol. 30, Part 2, 2002.
8 M. Heckele, W. K. Schomburg, "Review on micro molding of thermoplastic
polymers," JournalofMicromechanics andMicroengineering. Vol 14, 2004, p
Ri-R14
9 M. Rossi, I. Kallioniemi, "Micro-optical modules fabricated by high-precision
replication processes," Presentedat OSA topical meeting "Diffractiveoptics and
micro-optics" June 3-6, 2002; Tucson
10 Y. Xia, G. M. Whitesides, "Soft Lithography," Annual. Rev. Material.Science, 1998.
28:153-84
11 Yu-Chuan Su, Jatan Shah and Liwei Lin, "Implementation and analysis of polymeric
Microstructure replication by micro injection molding", J. Micromech. Microeng.
Vol. 14 No. 3, 2004 p 415-422
12 D. E. Hardt, B. Ganesan, Q. Wang, M. Dirckx, A. Rzepniewski, "Process Control in
Micro-Embossing: A Review." Singapore MIT Alliance Program in Innovation in
Manufacturing Systems Technology, Singapore, Jan. 2004
163
13 N. Roos, T. Luxbacher, T. Glinsner, K. Pfeiffer, H. Schulz, H. C. Sheer,
"Nanoimprint Lithography with a commercial 4 inch bond system for hot
embossing." Presented at SPIE's symposium Microlithography, Feb. 25-March 2,
2001
14 N. Roos, M. Wissen, T. Glinsner, H. C. Scheer, "Impact of vacuum environment on
the hot embossing process." Presented at SPIE's symposium Microlithography,
Feb. 22-28, 2003, Santa Clara CA
15 A. E. Bacon, S. Tiwari, L. Rathburn, "Nanoimprinting by hot embossing in polymer
substrates." National Nanofabrication Users Network, Cornell Nanofabrication
Facility, p 6-7.
16 L. Lin, Y. T. Cheng, C. J. Chiu, "Comparative study of hot embossed micro structures
fabricated by laboratory and commercial environments." Microsystem
Technologies. Vol 4, 1998, p 113-116.
17 B. Ganesan, Process controlfor micro embossing: Initial variabilitystudy. S.M.
Thesis, Massachusetts Institute of Technology, 2004.
18 D. Hardt, B. Ganesan, M. Dirckx, G. Shoji, K. Thaker, "Process variability in microembossing." Singapore MIT Alliance Program in Innovation in Manufacturing
Systems Technology, Singapore, Jan. 2005.
19 D. E. Hardt, "Manufacturing processes and process control." Reading from course
2.830, offered at the Massachusetts Institute of Technology Feb-May 2004.
Available via Open Courseware at: http://ocw.mit.edu/OcwWeb/MechanicalEngineering/2-830JSpring2004/Readings/index.htm
20 Obducat product catalog: http://www.obducat.com/pdf/ObducatProducts.pdf
21 Jenoptik products web page: http://www.jo-mikrotechnik.com/
22 EV Group products web page:
http://www.evgroup.com/markettoproduct.asp?MTPid=103
23 Suss MicroTec Brochure "Nanoimprinting Lithography."
24 L. Lin, C. J. Chiu, W. Bacher, M. Heckele, "Microfabrication using silicon mold
inserts and hot embossing." Presented at the Seventh international symbosium on
micro machine and human sciente, IEEE 1996.
25 F. P. Incropera, D. P. DeWitt, FundamentalsofHeat andMass Transfer, Fourth Ed.
John Wiley & Sons, 1996.
26 M. L. Boas, Mathematicalmethods in the physical sciences. John Wiley & Sons,
1983.
164
27 K. Ogata, Modern Control Engineering. Fourth Ed. Prentice Hall, 2002.
28 Melcor thermoelectric cooling products web page: http://www.melcor.com/tec.html.
29 Process News Magazine, Circulation heaters web page: http://www.processcontrols.com/HCS/hcs_recirc.html
30 Tranter PHE product web page for MaxChanger line:
http://www.tranterphe.com/phe/maxchanger/maxchanger.htm
31 Handbook ofHeat Transfer, Third Ed. W. M. Rohsenow, J. P. Hartnett, Y. I Cho,
Editors. McGraw-Hill, 1998.
32 A. Bejan, Convection heat transfer,Third Ed. John Wiley & Sons, 2004.
33 W. S. Janna, Design ofFluid Thermal Systems. PWS Publishing Co. 1998.
34 Spirex Sarco Learning Centre, section 6.3. Web site:
http://www.spiraxsarco.com/learn/default.asp?redirect=html/6_3_01.htm.
35 P. Skousen, Valve Handbook. McGraw-Hill, 1998.
165