Design of an
Ethanol Dehydration System
Jamie Hiltz Zack Taylor Mark Baier
Department of Chemical Engineering
University of Saskatchewan 2007‐2008
i
Abstract
Halo Consulting has been working on the design to purify ethanol with March
Consulting Associates Inc. The system was needed to dehydrate a liquid feed of 95 %
v/v ethanol – 5% v/v water to a minimum of 99.5 % v/v product. After Halo Consulting
examined a variety of different mass transfer applications to carry out this task, it was
decided that the use of pressure swing adsorption would be best.
The final design consists of two parallel columns one of which is adsorbing while the
other is either desorbing or in stand‐by. The liquid feed is heated and vaporized before
it is sent to the adsorbing column where it is dehydrated by a packed bed of Type 3A
molecular sieve. The regeneration time and adsorption times were found to be 6.33
and 23.9 minutes respectively. For the first 6.33 minutes of adsorption, 40 % of the dry
ethanol product is sent to help regenerate the desorbing column while the other 60 % is
condensed to the final product. For the remaining 17.61 minutes in the adsorption
cycle, the regenerating purge stream is no longer needed and 100 % of the product
stream is condensed to the final product. The physical properties of the columns and
the additional equipment were determined using mass balances.
ii
The costs of the equipment and the annual operating costs for this design were found.
An economic analysis was completed, comparing the cost of energy to break the
azeotrope with this design to that of azeotropic distillation and it was determined that
this design is more economically feasible.
A safety analysis was performed that consisted of a HAZOP analysis and a DOW Fire and
Explosion Index analysis on one of the columns. The most hazardous possible
deviations were determined to be high temperature and leaks. Recommended
preventative actions have been included in this report.
iii
Acknowledgments
The members of Halo Consulting would like to acknowledge the following people for their
direction and guidance throughout the duration of this design project:
Dr. Gordon A. Hill, Chemical Engineering 422 Advisor, Department of Chemical
Engineering, University of Saskatchewan
Dr. Hui Wang, Chemical Engineering 422 Advisor, Department of Chemical Engineering,
University of Saskatchewan
Dr. Richard Evitts, Faculty Advisor, Department of Chemical Engineering, University of
Saskatchewan
Dongmei Fang – Senior Process Engineer, March Consulting Associates Inc.
Tara Zrymiak – Senior Process Engineer, March Consulting Associates Inc.
iv
Table of Contents
1.0 INTRODUCTION 1 1.1 Background 1 1.2 Purpose and Proposed Design 1
2.0 LITERATURE SURVEY 2 2.1 Introduction 2 2.2 Distillation 3
2.2.1 Azeotropic Distillation 3 2.2.2 Pressure Swing Distillation 4
2.3 Thermal Swing Adsorption 5
3.0 DETAILED QUALITATIVE PROCESS DESCRIPTION 6 3.1 Introduction 6 3.2 Feed Preparation 7 3.3 Dehydration 8
3.3.1 Pressurization 10 3.3.2 Adsorption 10 3.3.3 Blow down 13 3.3.4 Regeneration 14
3.4 Cooling and Condensing 17 4.0 SIMULATION AND BALANCES 18
4.1 Introduction 18 4.2 Feed Preparation 19 4.3 Dehydration 21 4.4 Vacuum Pump 24 4.5 Condensation 25 4.6 Overall 26
v
5.0 EQUIPMENT DESCRIPTION AND SIZES 29
5.1 Introduction 29 5.2 Heat Exchangers 30
5.2.1 Heat Exchanger #1 30 5.2.2 Heat Exchanger #2 31 5.2.3 Heat Exchanger #3 31
5.3 Vacuum Pump 32 5.4 Adsorption Towers 33 5.5 Molecular Sieve 39 5.6 Valves and Piping 41
6.0 ECONOMICS 43 6.1 Introduction 43 6.2 Equipment Costs 44 6.3 Molecular Sieve Costs 46 6.4 Alternative Economic Comparison 46
7.0 SAFETY ANALYSIS 48
7.1 Introduction 48 7.2 Chemical Properties 49
7.2.1 Ethanol 49 7.2.2 Alumino Sillicate 50
7.3 Hazard and Operability analysis 50 7.3.1 HAZOP Strategy 50 7.3.2 HAZOP Conclusions 51
7.4 DOW Fire and Explosion Index Analysis 54 7.5 Process Safety Management 54
8.0 Conclusions 56 7.0 Recommendations 58 References 62
Appendix A: Process Flow Diagrams 59
vi
Appendix B: Mass Balances 66 Appendix C: Adsorption Data 70 Appendix D: Sizing Calculations 75
‐ Column Sizing 76 ‐ Breakthrough Curve Calculations 77 ‐ Equipment Sizing 81 ‐ Summary Tables 87
Appendix E: Economics Calculations 90 Appendix F: Piping and Instrumentation Diagram 94 Appendix G: Material Safety Data Sheets 96 Appendix H: HAZOP 111
vii
List of Tables
Table 4.1: Mole balance data for adsorption column with use of purge stream 22
Table 4.2: Mole balance data for adsorption column without use of purge stream 23
Table 4.3: Mole balance data for regeneration step 23
Table 4.4: Mole balance data for overall system 27
Table 5.1: Physical adsorption properties of adsorption column 36
Table 5.2: Typical properties of ZEOCHEM©. Z3‐03 38
Table 5.3: Ethanol dehydration equipment 42
Table 6.1: Summary of equipment economics 45
Table 6.2: Summary of alternative comparison 47
Table B.1: Mass balance for adsorbing column (Bed 1) 67
Table B.2: Mass balance for desorbing column (Bed 2) 68
Table B.3: Mass leaving system 69
Table B.4: Overall system mass balance 69
Table C.1: Table of given and calculated data for determining the breakthrough curve74
Table D.1: U‐Tube heat exchanger calculated data for pre‐heating feed 87
Tables D.2: Bayonet heat exchanger calculated data for feed vaporization 87
Table D.2a: Calculated data for sensible heating 87
Table D.2b: Calculated data for phase change 87
Table D.2c: Calculated data for superheating 88
viii
Tables D.3: Bayonet heat exchanger calculated data for product condensation 88
Table D.3a: Calculated Data for phase change 88
Table D.3b: Calculated Data for sensible cooling 88
Table D.4: Calculated Data for liquid ring vacuum pump 89
ix
List of Figures
Figure 3.1: Process flow diagram of entire system 9
Figure 3.2: Switching sequence for adsorption columns 9
Figure 3.3: Pressure swing adsorption cycle pressurization and adsorption of bed 11
Figure 3.4: Breakthrough curve 13
Figure 3.5a: Valve sequence for blow down step 14
Figure 3.5b: Valve sequence for regeneration step 14
Figure 3.6: Cycle steps for pressure swing adsorption 16
Figure 4.1a: Summary of data for U‐Tube heat exchanger to pre‐heat feed 20
Figure 4.1b: Summary of data for Bayonet heat exchanger to vaporize feed 20
Figure 4.2: Summary of data for adsorption column 21
Figure 4.3: HYSYS screen shot of vacuum pump 25
Figure 4.4: Summary of data for Bayonet heat exchanger to condense product 26
Figure 4.5: HYSYS screen shot of overall mass balance 27
Figure 5.1a: Theoretical curve 35
Figure 5.1b: Actual curve 35
Figure 5.2: Breathrough curves with varied bed heights 37
Figure 5.3: Visual of one adsorption column 40
Figure 5.4: Visual of alumino silicate 38
Figure A.1: Process flow diagram mimicking the dehydration system in HYSYS when Bed 2 (BAL‐2) is in regeneration 63
Figure A.2: Process flow diagram mimicking the dehydration system in HYSYS when Bed 2 (BAL 2) is done regenerating 64
x
Figure A.3: Process flow diagram of the ethanol dehydration system 65 Figure C.1: Isothermal data for water adsorption on a type 3A molecular sieve 71 Figure C.2: Water vapour isotherm at 120 for Type 3A molecular sieve 72 Figure C.3: Graph for the determination of equilibrium constant using Languir’s form 73 Figure F.1: Process and instrumentation diagram 95 Figure H.1: Summary of HAZOP analysis 112
xi
Roman Nomenclature
Symbol Name Units
Phase change area
Sensible heat area
Superheating area
Total area required for heat exchanger
Cross‐sectional area of the column
Annual cost $
Utility cost coefficient no units
Utility cost coefficient no units
Heat capacity of ethanol ·
Heat capacity of water ·
Heat capacity of mixed stream ·
Cost of fuel used to generate utility $
Concentration entering
adsorption column
xii
Concentration leaving
adsorption column
Column diameter
Diffusivity of ethanol into
water
Effective Diffusivity
Knudsen diffusivity
Particle diameter
Surface diffusivity
Pd Pore diameter
Pump down factor no units
Correction factor no units
G Gas superficial mass velocity ·
Heat of Adsorption ·
K Adsorption equilibrium constant no units
K Constant no units
Mass transfer coefficient
xiii
Column length
Bed length
Length of equilibrium zone
Length of unused bed
Molecular weight of ethanol molg
Molecular weight of water molg
Molecular Weight of the ethanol molg
and water mixture
Average Molecular Weight molg
Molecular weight of the mixed stream
constant no units
Mass flow rate
Mass flow rate entering
adsorption column
Mass flow rate leaving
adsorption column
xiv
Auxiliary plant capacity
Mass flow rate entering
vacuum pump
Reynolds number no units
Schmidt number no units
Sherwood number no units
Pressure entering
adsorption column
Pressure entering vacuum pump
Pressure leaving
adsorption column
Pressure leaving vacuum pump
Pressure of the purge stream
Plant cost index no units
Heat duty
Volumetric flow rate entering
adsorption column
xv
Volumetric flow rate leaving
adsorption column
Adsorption capacity ·
Gas constant · or
·
Pump capacity
Surface area of molecular sieve
Time required to reach a specific
vacuum level
Temperature entering
adsorption column
Temperature leaving
adsorption column
Temperature of the purge stream
Adsorption time
Breakthrough time
Thickness of the column
Regeneration time
Ideal adsorption time for
Vertical breakthrough
xvi
Overall heat transfer coefficient · ·
Utility Price $
Superficial velocity
Maximum superficial velocity
Volume of column to be evacuated
Volume of adsorbent
Volume of the column
Volume if void in the column
Molar flow rate entering
adsorption column
Molar flow rate leaving
adsorption column
Shaft work
xvii
Greek Nomenclature
Intrinsic efficiency no units
Particle porosity no units
Latent heat of ethanol
Latent heat of water
Latent heat of vaporization
of the mixed stream
Viscosity entering · adsorption
column
Viscosity leaving ·
adsorption column
Atomic diffusion volume of Carbon no units
Atomic diffusion volume of Hydrogen no units
Atomic diffusion volume of Oxygen no units
Average molecular velocity
∑ Diffusion volume of ethanol no units
∑ Diffusion volume of water no units
Porosity no units
xviii
Bulk density
Density of fluid entering
adsorption column
Density of fluid leaving
adsorption column
Particle density
Other
∆ Log mean temperature
1
Chapter 1.0: Introduction
1.1 Background
For the past eight months, Halo Consulting has been working on a design project for
March Consulting Associates Inc. March Consulting was incorporated in 1999 and their
main office building is located in Saskatoon, Saskatchewan. They offer project and
design services for many different disciplines of engineering, and always strongly stress
energy conservation.
1.2 Purpose and Proposed Design
The purpose of this project was to design a dehydration system that would purify a feed
of 95 % v/v ethanol – 5 %v/v water to a minimum of 99.5 % v/v ethanol product. It was
proposed by March Consulting that two columns be designed and sized to use pressure
swing adsorption with a Type 3A molecular sieve.
2
Chapter 2.0: Literature Survey: Alternative Processes
2.1 Introduction
Upon performing the initial research, a vast assortment of directions for this design
were discovered and considered. This collection of alternative processes was carefully
analysed in order to depict the design which was most favourable in regards to the
quality of the product, economic considerations, and energy consumption. The first and
foremost necessary resolution was to decide which mass transfer application was to be
used to dehydrate the ethanol. The three modes of mass transfer investigated were
distillation, thermal swing adsorption, and pressure swing adsorption.
3
2.2 Distillation
Distillation is a very common method of liquid‐liquid separation that works by “the
application and removal of heat to exploit differences in relative volatility” (Africa
1996). Applying heat to a mixture allows its components with lower boiling points to
vaporize into a gas phase that travels to the top of the column, separating it from the
components with higher boiling points because they retain a liquid phase and travel to
the bottom of the column. Because this method of separation is based on boiling
points, it is difficult to separate azeotropic feeds. An azeotropic feed is “a liquid mixture
that maintains a constant boiling point and that produces a vapour of the same
composition as the mixture” (Africa 1996). In other words, because the inlet feed in this
design has a 95% v/v– 5% v/v water, it is classified as an azeotropic feed and could not
be easily separated by the use of simple distillation.
2.2.1 Azeotropic Distillation
Azeotropic distillation can be successfully used to break the azeotropic feed but
requires certain additives to do so. It can be classified into two types; homogeneous
azeotropic distillation and heterogeneous azeotropic distillation.
Homogeneous azeotropic distillation requires the use of an entrainer which is a,
“separating agent that forms an azeotrope with one of the components of the binary
feed” (Africa 1996). The new azeotrope is easier to separate than the first one and is
4
sent to a second column that is operating at an appropriate pressure to break it. This
results in the formation of many azeotropes in the attempt to break the initial one. This
makes the design and simulation very difficult as the behaviours of these azeotropes are
very unpredictable. The entrainers that are required are also very expensive.
Heterogeneous azeotropic distillation has the ability to separate azeotropic feeds while
using less entrainment than homogeneous azeotropic distillation with a self‐entraining
system, but the design and simulation is still very difficult and it results in large recycle
rates.
2.2.2 Pressure Swing Distillation
Pressure swing distillation is a specialized type of distillation that has the ability to
separate azeotropic feeds without the aid of additives. This process has a series of
distillation columns that operate at different pressures in order to break the azeotrope.
The first column operates at one pressure to separate a small amount of ethanol from
the mixture. The distillate from this column is then sent to another distillation column
that operates at a different pressure, breaking the azeotrope, and separating a little
more ethanol. The bottoms of the column is sent back to the feed as a recycle stream.
This repeats with a number of towers until the desired volume percentage of ethanol is
achieved. Although successful, this method is not feasible for the purification of
5
ethanol to 99.5% v/v as it requires high energy consumption. It also requires many
large columns which would result in high capitol costs. “Pressure swing distillation can
be used to break an ethanol‐water mixture that forms an azeotope. The process
consists of three or more columns operating at different pressures” (Africa 1996).
2.3 Thermal Swing Adsorption
Thermal swing adsorption, also known as temperature swing adsorption, is an effective
way to remove impurities from gas streams. The problem with thermal swing
adsorption is that in order to achieve a temperature range that sufficiently affects the
separation desired, a long swing time is required. From this arises the need for larger
equipment and large amounts of energy consumption, resulting in high operation and
maintenance costs. These costs can be reduced by enhancing the system with a
“microchannel architecture” (Africa 1996), improving mass transfer and therefore
allowing smaller sized equipment and shorter cycles. The cost of enhancing the system,
however, still leaves this process as not economically feasible.
It was decided that the best method of mass transfer for this problem was to use
pressure swing adsorption.
6
Chapter 3.0: Detailed Qualitative Process Description
3.1 Introduction
The feed into the dehydration system is a two component liquid mixture of water and
ethanol. This mixture is coming in at a liquid volume percent of 95 % ethanol and 5%
water, which is regarded as an azeotropic mixture. To obtain a liquid product of
minimum 99.5 % v/v ethanol the feed must go through three main steps: (1) pre‐
heating and vaporizing of liquid feed, (2) dehydration of wet ethanol gas by pressure
swing adsorption, and (3) cooling and condensing of dry ethanol gas to a liquid. A
process flow diagram (PFD) for the proposed dehydration system can be seen in Figure
3.1.
7
3.2 Feed Preparation
Before the azeotropic liquid feed mixture can be sent through the columns, certain
preparation steps need to be carried out. The adsorption is being performed by a type
3A alumino silicate that can be damaged by the presence of liquid. Therefore, the feed
stream to the first column needs to be vaporized and brought to a temperature and
pressure that will eradicate any possibility of condensation within the columns.
The 95 % v/v ethanol feed is first pre‐heated using a U‐tube heat exchanger. The liquid
feed passes through the shell side of the heat exchanger where it is heated with hot
product gas that is passing through the tube side of the heat exchanger. No
vaporization of the feed will occur in this first heat exchanger, its only purpose is to
raise the temperature of the liquid to prepare it for vaporization. The hot liquid mixture
is then passed through a bayonet heat exchanger where it is completely vaporized. This
is completed by passing the hot liquid through the shell side where it is vaporized by
superheated steam that is passing through the tube side. The resulting gas will be at a
temperature and pressure where the chance of condensation within the column is
eliminated. This is particularly important because the formation of liquid within the
column could damage the molecular sieve and reduce its adsorption capacity.
8
3.3 Dehydration
Once the feed has been prepared, it can be sent to the adsorption column to be
dehydrated. The hot wet ethanol gas is passed through one of two vertical adsorption
columns that are aligned parallel to each other. This can be seen in Figure 3.3. These
two columns utilize Pressure Swing Adsorption (PSA) to dehydrate the ethanol. In PSA,
each column, “operates alternately in two half‐cycles of equal duration” (Henley 2006).
In this design, a whole cycle involves:
1) Pressurization of the column with the feed;
2) Adsorption at elevated pressure;
3) Depressurization by vacuum pump;
4) Desorption by purge stream at lower operating pressure;
5) Stand‐by
This cycle is a slightly modified form of the Skarstrom cycle where the cycle entails, “(1)
pressurization followed by adsorption, and (2) depressurization (blow down) followed
by a purge” (Henley 2006). The simple set up of columns and valves in Figure 3.1 shows
the flow pattern in this subsystem of the design throughout a complete cycle. At the
beginning of cycle, the saturated bed will undergo a regeneration process to desorb the
water from the bed while the other uses adsorption to dehydrate the ethanol and
becomes saturated. The adsorbing and desorbing columns are represented by Bed 1
and Bed 2, respectively in Figure 3.3. The following will describe the cycle for Bed 1.
9
Figure 3.1 Process flow diagram of entire system.
Figure 3.2 Switching sequence for adsorption columns
10
3.3.1 Pressurization
Initially, Bed 1 will exist at a pressure of 101.3 kPa. When the cycle starts, valves 1 and
7 open rapidly, allowing flow of wet ethanol gas to go through Bed 1. The wet ethanol
gas flows into the top of Bed 1 at an elevated temperature and pressure. This raises the
pressure of the column to match that of the feed stream. The pressurization step takes
about 10 seconds which is relatively short in comparison to the adsorption and
regeneration steps. During this period, adsorption of water onto the molecular sieve is
initiated, but the majority of adsorption will take place in the next step.
3.3.2 Adsorption
Once Bed 1 has been pressurized, a pressure transmitter on the column signals a
control switch that allows the adsorption step to begin. This step starts with the
opening of valves 5 and 6, which allows the regeneration gas to flow into Bed 2. With
valves 1, 4, 7, 5, and 6 open, a continuous flow of gas will be allowed through both of
the beds. This can be seen below in Figure 3.3. Wet ethanol gas, entering at the same
operating conditions as in the first step, will be continuously dehydrated in Bed 1,
resulting in a dry ethanol gas product flowing out the bottom.
11
Figure 3.3: Pressure‐swing adsorption cycle‐pressurization and adsorption of bed 1
This dry ethanol is then split into two streams of 40 and 60 %. The 40 % is sent to Bed 2
where it is used as a regeneration gas. The remaining 60% is to be sent for further
processing. Splitting the dry ethanol stream is controlled by valves 6 and 7, which will
be partially open to allow/restrict the amount of product into Bed 2.
The concentration of water in the outlet product will initially be zero because essentially
all of the water in the wet ethanol is removed. The assumption that complete
separation of water from ethanol would occur was based on theoretical data and made
for calculation purposes. In reality, complete separation would not occur. As time
passes, the amount of molecular sieve saturated with water increases, reducing the
capacity of the bed. Eventually, small amounts of water will been seen in the product.
The time at which water is first observed in the product is referred to as the
breakthrough time. If the column is allowed to operate past the breakthrough time,
12
increasing concentrations of water will be observed in the product until a point where
the concentration of water in the product is equal to that of the feed. At this time the
molecular sieve would be considered totally saturated and would have no capability to
adsorb water. In this design, the column is allowed to operate past the breakthrough
time until the product has a concentration of 99.5% v/v ethanol. The time at which a
99.5% v/v ethanol is reached was calculated using theoretical breakthrough curve.
Using the breakthrough curve in Figure 3.4, a time of 23.9 minutes was found to
correspond to a purity of 99.5% v/v ethanol and therefore marked the end of the
adsorption step for Bed 1. A detailed derivation of the systems breakthrough curve and
determination of cycle times will be covered in chapter 5.
The regeneration of Bed 2 takes less time than the adsorption of Bed 1, and therefore
the regeneration gas is not needed for the entire 23.9 minutes. A set of controls has
been implemented to measure the concentration of water in the stream leaving Bed 2.
Once the concentration of water in this stream is the same as the inlet feed to Bed 2,
the controls signal the splitters to discontinue the purge stream. At this time, the feed
rate to the entire system is also decreased. This was calculated to occur at
approximately 6.33 minutes. The feed rate is decreased to allow for a constant product
flow. The majority of the design calculations were performed using values obtained
13
from simulation with the purge stream was still in use. This was done so that a steady
state assumption could be made.
Figure 3.4: Breakthrough curve
3.3.3 Blow down
The next step in the PSA cycle is the blowing down of Bed 1. It is necessary because the
column must be depressurized in order for the regeneration step to occur. This step
begins with the rapid opening of valves 2, 3, and 8 and the closing of valves 1, 4, 5, 6,
and 7. This is shown in Figure 3.5a. Once the valves have switched, the liquid ring
vacuum pump will automatically start. The purpose of this pump is to bring the column
to a pressure that is below atmospheric as a preparation step for the regeneration
0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1
0 10 20 30 40 50 60
c/cf
Time, min
14
process. This step is comparable to that of pressurization because the duration of this
step is relatively short and marks the beginning of the regeneration step. The vacuum
pump will run continuously for the entire duration of the regeneration step.
3.3.4 Regeneration
Regeneration is the last step in the PSA cycle and is done to remove the adsorbed water
from the molecular sieve bed so that it may be reused in the next cycle. Pressure
sensors and controls have been added to both columns to aid in switching the beds
from adsorption to desorption. Once the pressure sensor on Bed 1 measures that a low
enough pressure has been achieved, valves 5 and 6 open, beginning the regeneration
step displayed in Figure 3.5b. The opening of these valves allows dry ethanol gas to
enter the bottom of Bed 1. Valve 5 will be partially closed, throttling the regeneration
gas to the same pressure as the column. The low operating pressure given by the
running vacuum pump allows the water to be easily removed from the molecular sieve
by the dry ethanol gas entering the bottom.
15
Figure 3.5a Valve sequence for blow down step Figure 3.5b Valve sequence for regeneration step
The amount of purge gas and pressure within the bed determine the amount of time
needed for regeneration. A large amount of purge gas and a lower column pressure
results in a shorter regeneration time. The regeneration step for this molecular sieve is
about 6.33 minutes. This is acceptable as it is much less than the time needed for
adsorption, eliminating any possible downtime. This time is often determined using
experimental data, but this estimation was the result of a correlation involving the
pressures and flow rates of the feed and purge streams. A concentration sensor at the
top of the column will shut off the vacuum pump and close valves 5 and 6 when no
more water is observed in the product, as described above.
The complete removal of water from the molecular sieve is very difficult and is
considered uneconomical because equipment and operating costs would be very high.
The calculations for this design were performed assuming that complete removal of
16
water was achieved because finding the exact amount water that can be removed at
these conditions would require experimental data that was not available.
With the regeneration time being less then the adsorption time there will be a period
where one of the towers is not in operation and is standing by. This is more desirable
then having a regeneration time greater then the adsorption time, which would result
in a semi‐continuous flow of product. To remove the problem of standby, additional
adsorption columns could be introduced into the system. A complete sequence of cycle
steps in PSA is shown in Figure 3.6
Figure 3.6 Cycle steps for pressure swing adsorption (Ruthven 1994)
17
3.4 Cooling and Condensing
The remaining 60 % of the dry ethanol gas that is not used for the regeneration step is
sent through two heat exchangers where it is cooled to the final liquid product. The hot
dry gas is first passed through a U‐tube heat exchanger where it is cooled by the liquid
feed. This is the same heat exchanger that was used for pre heating the liquid feed.
The cooled gas is finally passed through a bayonet heat exchanger where it is
condensed using cooling water. The cooled gas will condense as it passes through the
tubes, resulting in a final liquid product of 99.5 % v/v ethanol.
18
Chapter 4.0: Mass and Energy Balances: Simulations
4.1 Introduction
The simulation package HYSYS was used for simulating the vaporizing and condensing
processes within the design. The adsorption and desorption processes could not be
modelled in HYSYS as it is not fit for adsorption processes. They were therefore
modelled by hand using Excel. A total system mass balance was done with HYSYS but an
energy balance for the system could not be completed because the energy transferred
in the adsorption process could not be simulated.
19
4.2 Feed Preparation
The 95 % v/v liquid ethanol enters the system at an initial molar rate of 41.86 ·
until the regeneration of the desorbing bed is complete. Once the regeneration is
complete, the feed rate is decreased to 25.11 · since no more dry gas is needed
for purging and a constant liquid product flow rate of 1250 is desired.
For both of the flow rates, the temperature and pressure of the feed entering the U‐
tube heat exchange is 35 and 129 kPa. This feed is heated to a temperature and
pressure of 85˚C and 132 kPa by the hot, dry ethanol gas leaving the adsorption column
at 120˚C and 175 kPa. The hot gas, flowing at a constant rate of 21.84 ·
, allows the
liquid feed to be heated but does not vaporize it. The hot liquid feed is then completely
vaporized to 120˚C and 250 kPa using superheated steam in a bayonet heat exchanger.
The superheated steam flow rate will adjust accordingly with the inlet flow rate so that
this temperature and pressure is reached and held constant before entering the
adsorption column. Flow rates, pressures, temperatures, and duties for the stream and
heat exchangers, for the first 6.33 minutes, were found using the HYSYS simulation and
are summarized in Figure 4.1.
20
Figure 4.1a Summary of data for U‐Tube heat exchanger to pre‐heat feed
Figure 4.1b Summary of data for Bayonet heat exchanger to vaporize feed
21
4.3 Dehydration
The wet ethanol gas is then passed into the first heat exchanger where adsorption will
take place at 120˚C and 250 kPa. A mole balance was first completed around
adsorption column for the 6.33 minute period where 40 % of its product is being used
to regenerate the desorbing column. An adsorption rate of 5.47kg‐mole/h resulted in
0.577 kg‐mole of water being adsorbed during this time. The adsorption of water is
slightly exothermic as the molecular sieve reacts with water to produce heat.
Approximately 2424 kJ of heat is given off during this period. Since the cycle times of
PSA are so short the columns are assumed to be isothermal. The flow rates,
temperatures, pressures and component mole fractions for the adsorption column, Bed
1, can be seen in Figure 4.2 and Table 4.1.
Adsorption Column
Figure 4.2 Summary of data for adsorption column
T=120˚C
P=250 kPa
T=120˚C
P=175 kPa
22
Table 4.1 Mole balance data for adsorption column with use of purge stream
Streams Mole Fraction Flowrate (kgmole/h)
Water Ethanol Feed Water Ethanol IN 0.14 0.86 41.86 6.04 35.81 OUT 0.02 0.98 36.39 0.58 35.81 ACCUMULATED 1 0 5.47 5.47 0.00
The outlet pressure of 175 kPa was determined using Ergun’s equation for pressure
drop in a randomly packed column. The pressure drop was found to be around 75 kPa
and does not cause damage to the molecular sieve.
A second mass balance was completed around the adsorption column for the 17.61
minutes that the purge stream is no longer needed and the entire product from the
adsorption column is sent to be condensed to the final liquid product. For this period
there is a water adsorption rate of 3.27 ·
, which results in 0.96 kg‐mole of water
being adsorbed and 4041 kJ of heat being produced. Values for the component mole
fractions and flow rates are shown in Table 4.2.
23
Table 4.2 Mole balance data for adsorption column without use of purge stream
Streams Mole Fraction Flowrate (kgmole/h)
Water Ethanol Feed Water Ethanol IN 0.14 0.86 25.11 3.63 21.48 OUT 0.02 0.98 21.84 0.35 21.48 ACCUMULATED 1 0 3.27 3.28 0.00
For the total adsorption time, 23.9 minutes, a total of 1.54 kg‐mole of water is adsorbed
onto the molecular sieve. This amount of water will essentially be the amount that
needs to be removed from the desorbing column. The water will be desorbed at a rate
of 14.81 ·
and will act as the generation term in the mass balance equation for that
column. Table 4.3 shows the component mole fractions and flow rates for removing
the 1.54 kg‐mole of water during the regeneration step (6.33 minutes).
Table 4.3 Mole balance data for regeneration step
Streams Mole Fraction Flowrate (kgmole/h)
Water Ethanol Feed Water Ethanol IN 0.02 0.98 14.56 0.23 14.33 GENERATED 1.00 0.00 14.58 14.58 0.00 OUT 0.51 0.49 29.14 14.81 14.33
The outlet product of the regenerating column is considered a by‐product of the
process and is sent out of this design to be further processes.
24
4.4 Vacuum Pump
The vacuum pump was modelled as a compressor using HYSYS. This vacuum pump is of
a liquid ring type and its purpose is to maintain a pressure of 50kPa in the desorbing
bed. This pump takes the wet desorbed gas and sends it away to be further processed.
The gas entering the vacuum pump is flowing at a rate of 29.14 ·
, which is equal to
the outlet flow rate. The exiting pressure of the gas will be at 101.3 kPa. HYSYS
calculated the pump to have a power rating of 26.10 kW with an efficiency of 75%. This
power rating is very close to power rating of 21.73 kW that was completed by hand
calculations for the specific pump. The pump is continuously running during the
regeneration period and is shut off only when the bed is fully desorbed, after 6.33
minutes. A HYSYS screen shot of the vacuum pump can be seen in Figure 4.3 along with
the corresponding values for the streams and the vacuum pump. The purge gas
entering the desorbing bed is first throttled down to a pressure of 50 kPa by the
butterfly valve V‐11 in Figure 3.1.
25
Figure 4.3 HYSYS screen shot of vacuum pump
4.5 Condensation
The dry ethanol gas, that is sent to be condensed, flows into the U‐tube heat exchanger
at a constant flow rate of 21.84 ·
. The product gas in stream 5, is at a temperature
and pressure of 120˚C and 175 kPa, and is cooled to 91 ˚C and 165 kPa by the liquid feed
stream 1. At this point some condensation of the dry ethanol gas has already started to
occur. Figure 4.1 is a representation of the U‐tube heat exchanger. The remainder of
the gas is completely condensed to 30˚C and 130 kPa by cooling water in a bayonet heat
exchanger. The final 99.5% v/v ethanol product is flowing out of the system at 21.85
· or 1250 which was specified by March Consulting. Figure 4.4 is a representation
of the bayonet heat exchanger.
26
Figure 4.4 Summary of data for Bayonet heat exchanger to condense product
4.6 Overall
After mass balances around the two columns were completed, an overall mass balance
for the system for a half cycle was done. The overall mass balance for the system was
slightly more complicated due to the fact that a certain amount of water is being
accumulated in the adsorption bed and being desorbed in the regenerating bed. There
is also a dynamic step change that occurs in inlet flow rate when one of the beds is done
regenerating. This also complicates the mass balance.
It can be seen in Table 4.4 that after the adsorption for one bed is completed, the
amount entering the system is equal to the amount leaving of the system.
27
Table 4.4 Mole balance data for overall system
Stream Substance (kgmole)
Water Ethanol Total IN (feed) 1.70 10.08 11.79 ‐ OUT (Product) 0.138 8.58 8.71 ‐ OUT (By‐Product) 1.56 1.51 3.08 = 0.00 0.00 0.00
It should be noted that the mass balance was completed while maintaining the final
liquid product at 99.5% v/v ethanol. In reality, this composition will vary between 100
and 99 % v/v ethanol. Figure 4.5 shows how HYSYS was used to simulate an overall
mass balance for the time period 6.33 minutes.
Figure 4.5 HYSYS screen shot of overall mass balance
The dashed blue line represents the amount of water that is accumulated in the
molecular sieve over a half cycle and needs to be removed in the regeneration step.
28
The 5.47 ·
is the rate of water that is adsorbed in the first 6.33 minutes, and the
9.11 ·
is the rate at which the water was adsorbed in the time period of 17.61
minutes. Complete stream and unit operation info for both mass and energy balances
can be found in the DESIGN.HSC file for HYSYS simulation in the attached CD.
29
Chapter 5.0: Equipment Description and Sizes
5.1 Introduction
Once the mass balances were completed, each piece of equipment in the process could
then be sized. The equipment sizes were calculated using the “Short‐Cut Equipment
Design Method” developed by Ulrich and Vasudevan in their textbook, Chemical
Engineering Process Design and Economics: A Practical Guide. The final design, as
mentioned earlier, consisted of three heat exchangers, one vacuum pump, and two
identical towers. Since ethanol does not have any corrosive properties and the
pressures and temperatures within the system are not to extreme, all of the equipment
for this design will be constructed out of carbon steel. Thought was invested into using
some stainless steel within the system, however, this would be an unnecessary expense
and introducing a second type of metal brings accelerated bi‐metallic corrosion into the
system. A summary of the equipment sizes can be seen in Table5.3 at the end of this
chapter.
30
5.2 Heat Exchangers
5.2.1 Heat Exchanger #1
In this design there were three heat exchangers used, all of which are of shell and tube
type. The first heat exchanger is used to pre‐heat the feed stream and cool the final
product stream. Using Table 4.12 from Ulrich’s textbook, it was determined that a U‐
Tube heat exchanger had the compatibility and service ratings that were appropriate for
this task and was therefore chosen.
The U‐tube heat exchanger uses the product stream from the adsorption column as the
heating source to pre‐heat the liquid feed at cool the product simultaneously. The final
area of this heat exchanger was found to be 2.21m2. Guidelines given by Ulrich were
used to determine which stream would flow through the shell side and which would
flow through the tube side. Because it is flowing at a higher pressure, it was
determined that the hot product stream should flow through the tube side. The cool
feed stream would therefore flow through the shell side.
31
5.2.2 Heat Exchanger #2
The second heat exchanger takes the pre‐heated feed from heat exchanger #1 and
vaporizes it using superheated steam as a heating source. Because of the vaporization
and superheating involved, a U‐tube heat exchanger could not be used. Of the five
types of heat exchangers given in Table 4.12 (Ulrich 2004), it was found that a bayonet
heat exchanger is the only one able to handle the superheating aspects of these
conditions. The final area of this heat exchanger was found to be 15.22m2 and the
steam would flow through the tube side, while the feed would flow through the shell
side.
5.2.3 Heat Exchanger #3
The third heat exchanger is used to cool and condense the semi‐cooled product stream
coming out of heat exchanger #1. Because there is condensation involved, it was
determined that another bayonet heat exchanger would serve best for this task. The
final area of this heat exchanger was found to be 10.02m2 and the semi‐cooled product
would flow through the tube side, while the cooling water would flow through the shell
side.
32
5.3 Vacuum Pump
The vacuum pump chosen for our system was a Liquid Ring Vacuum Pump as, “Liquid
ring vacuum pumps are commonly used to handle “wet” gas mixtures” (Aliasso 2003).
In these types of mixtures, it is possible and common for the lubricant within the pump
to be washed away, causing “premature failure” (Aliasso 2003). Rather than using a
lubricant, liquid ring vacuum pumps eliminate any metal on metal contact to reduce the
chance of failure. This also reduces the wear and tear on the pump, extending its
lifespan. An impeller inside the pump rotates and throws the water by centrifugal
force, creating a liquid ring with the casing and thus generates compression. As the gas
enters, it is entrapped by the impeller blades and the liquid ring. As the impeller
rotates, it creates a compression on the gas and forces it though the pump outlet. To
eliminate the possibility of contamination from the liquid to the gas phase, the pump
will use water to create the liquid ring.
The equation used to determine the pump size was
TFVS ×
= (5.1) (Graco 2008).
where S is the vacuum pump size in cubic feet per minute, V is the volume of the
column to be evacuated in cubic feet, F is the pump down factor, and T is the time
33
required to reach a specific vacuum level in minutes. Using this equation, along with
the values, V =15.9 ft3, F=2, and T=0.17 min, and some unit conversions, it was
determined that the pump had a maximum pumping capacity of 325 and a shaft
power of 18 kW. This pump was used to bring the pressure of the column down to
approximately 50 kPa. The final output pressure on the pump is at atmospheric
pressure (101.3kPa).
5.4 Adsorption Columns
The sizing of the adsorption column was considered the most important aspect of the
design. In most instances the design of an adsorption column is based on experimental
data, but for this design this was not available. The sizing of the adsorption bed was
completed using equations and guidelines given by Henley and ZEOCHEM, the
manufacturer of the molecular sieve being used in designed adsorption column.
The first step in designing an adsorption column is to calculate the diameter of the
molecular sieve bed. The diameter of the bed should be as small as possible without
damaging the molecular sieve, “A diameter small enough to maintain a turbulent gas
flow is necessary, otherwise the mass transfer characteristics are very poor because of
the increased film resistance to mass and energy transfer”(ZEOCHEM 2007‐2008). The
34
maximum gas velocity before erosion or crushing of 1/8” beads is approximated by the
following equation given by ZEOCHEM:
, 61.5/ (5.2)(ZEOCHEM 2007‐2008).
For a downward maximum velocity of 0.697 , a bed diameter for the adsorption
column was calculated to be 0.55 m.
Once the bed diameter was calculated, the amount of molecular sieve needed for water
adsorption could be determined. Since no experimental data was available, a constant
pattern scale up of an adsorption column could not be readily completed. Therefore,
estimations of bed heights were made for an ideal fixed‐bed adsorption column and
later used to model a real fixed‐bed adsorber. An initial bed height was determined by
calculating the length of the equilibrium zone (LES), which was completed by doing a
water mass balance around the system. The amount of water adsorbed in each cycle
was found by multiplying the adsorption rate by the duration of the adsorption step. A
typical cycle time for PSA ranges from 5 ‐ 30 minutes, so for the purpose of this design
the adsorption time was taken as two‐thirds of maximum value, 20 minutes. In one
cycle, 1.82 kg‐mol of water is adsorbed. Based on this value, the amount of molecular
sieve was predicted using the equilibrium loading capacity for the molecular sieve found
in Table 5.2. ZEOCHEM recommends using half of this value, 10% ·
adsorbent,
35
because “polar compounds in the gas will reduce the equilibrium capacity” (ZEOCHEM
2007‐2008). A resulting LES of 2 m is required for the bed.
This length is not acceptable because there are many assumptions associated with it,
especially that, “local equilibrium between the fluid and the adsorbent is achieved
instantaneously, resulting in a shock like wave, called a stoichiometric front, that moves
as a sharp concentration front through the bed” (Henley 2006). This type of wave is
shown in Figure 5.1a. In a real fixed bed adsorber this will not occur and there would be
a mass transfer zone (MTZ) and a length of unused bed (LUB) with a concentration front
like that in Figure 5.1b. This 2 m length is an initial guess in modelling a breakthrough
curve that will be used in determining a true bed length.
Figure 5.1a Theoretical curve Figure 5.1b Actual curve (Henley 2006).
36
A breakthrough curve is a relationship between the ratio of the product and feed water
concentrations, (c/cF) versus time. Using this curve the time that corresponds to a c/cF
of 0.068 (max ratio giving 99.5% v/v ethanol) can be found. A theoretical breakthrough
curve for the system was approximated by Klinkenberg’s equation:
1 erf √√
(5.3)
Where,
(5.4)
And,
(5.5)
The adsorption equilibrium constant, K, was determined by fitting equilibrium
absorption data given by ZEOCHEM to a linear adsorption isotherm of the form q=Kc.
The overall mass transfer co‐efficient, k, was then determined by methods developed in
Henley and Seader’s, Separation Process Principles, which involves the relationship:
(5.6) (Henley 2006).
The values of K and k for the adsorption bed were found to be 245 and 0.309 s‐1
respectively. The values used in the development of the theoretical breakthrough curve
can be found in Table C.1. The height was then scaled up until a breakthrough time
37
around 20 minutes was achieved. The resulting height of 3.5 m was chosen, giving an
adsorption time of around 24 minutes. The resulting breakthrough curves are shown in
Figure 5.2
Figure 5.2 Breakthrough curves with varied bed heights
Once the final dimensions of the bed determined, the pressure drop was verified using
Eurgan’s equation. A pressure drop of 75 kPa was found across the bed which is an
acceptable pressure drop.
The column is to be constructed from carbon steel as corrosion effects are not a
problem in this process. Using correlations from Ulrich’s, the wall thickness for a
vertical cylindrical column was found to be 4 mm. All physical adsorption column
0
0.2
0.4
0.6
0.8
1
1.2
0 10 20 30 40 50
c/c F
Time (min)
2 meter bed
3 meter bed
3.5 meter bed
38
properties can be found in Table 5.1. A visual of an adsorption column can be seen in
Figure 5.3.
Figure 5.3 Visual of one adsorption column
Table 5.1 Physical adsorption properties of adsorption column
unit value
Material Carbon Steel
Column inner diameter m 0.55
Column wall thickness mm 4
Column height m 4.5
Bed height m 3.5
39
5.5 Molecular Sieve
The adsorbent used within the two identical towers is a type 3A molecular sieve zeolite.
This is the most commonly used adsorbent in the application of dehydrating ethanol.
This molecular sieve, “is the potassium form of the A‐type structure; and has an
effective pore opening of 3 Angstrom (0.32 nm)” (ZEOCHEM 2007‐2008). In Type A
structured zeolite, “the tetrahedral are grouped to form a truncated octahedron with a
silica alumina tetrahedron at each point” (ZEOCHEM 2007‐2008). A visual of this can be
seen in Figure 5.4. The structure is represented by the chemical formula:
0.45 K2O ∙ 0.55 Na2O ∙ Al2O3 ∙ 2 SiO2 ∙ XH2O (ZEOCHEM 2007‐2008).
For the design calculations, values given by ZEOCHEM for their ZEOCHEM® Z3‐03
were used. These values can be seen in Table 5.2. ZEOCHEM is large manufacturer of
commercially used adsorbents and, “supplies around 80% of the world market for
ethanol dehydration“(ZEOCHEM 2007‐2008). Their ZEOCHEM® Z3‐03 is specifically
designed for drying ethanol using pressure swing adsorption.
This molecular‐sieve zeolite is polar‐hydrophilic in nature and therefore strongly
adsorbs water. In the case of this design adsorption of water onto this adsorbent will
purely physical. In physical adsorption, “the intermolecular attractive forces between
molecules of a solid and the gas are greater than those between molecules of the gas
40
itself” (Henley 2006). These forces are commonly referred to as van der Waals
interactions.
Figure 5.4 Visual of alumino sillicate
The physical adsorption of just water is directly related to the effective pore size of the
molecular sieve. This adsorbent will absorb any molecule with a diameter less then 0.32
nm and exclude those that are larger. A water molecule has a diameter of 2.8 Angstrom
and is effectively adsorbed onto the molecular, whereas ethanol has a diameter of 4.4
Angstrom and is therefore excluded.
Table 5.2 Typical properties of ZEOCHEM® Z3‐03
unit valueTapped bulk density, EN ISO 787‐11 kg/m3 750 Bead size nominal mm 2.5‐5 Crush Strength N 70 Equilibrium water adsorption capacity, @20°C/50%rh/24h % 20 Residual water content , 550°C as shipped % 1.5 Heat of adsorption kJ/kg water 4200 Specific heat (approx) kJ/kg°C 1.07 (ZEOCHEM 2007‐2008).
41
5.5 Valves and Piping
The valves chosen for our design were pneumatic butterfly valves. The rationale for
choosing butterfly valves was that they are “one of the most successful high‐
performance valves” (Dickenson 1999). Dickenson also mentions that these valves are
ideal for smaller units, which reduces cost, weight and space requirements, and that
they have few parts, creating easy maintenance, installation, and operation. They were
also found to have a high rating for both liquid and gas services, on/off switching,
throttling, flow control, and quick opening. The purpose of having them pneumatically
operated is so that a control board operator can easily monitor and manipulate the
position that the valve is in. The piping being used in this design is once again carbon
steel with a inner diameter of approximately 6 inch and a thickness of inch. This was
determined using Perry’s Handbook.
42
Table 5.3 Ethanol dehydration equipment
Equipment Type Equipment Name Size
Heat Exchangers
U‐Tube E‐112 1.96 m2
Bayonet (Heating) E‐122 15.2 m2
Bayonet (Cooling) E‐132 10.0 m2
Vacuum Pump Liquid Ring G‐113 21.7 kW
Adsorption Tower Tower 1 or 2
D‐110
D‐120
Diameter: 0.55 m
Height: 4.5 m
*Sample calculations can be seen in Appendix D
43
Chapter 6.0: Economics
6.1 Introduction
Once sized, the prices of each piece of equipment were then estimated using the
program EconExpert (EconExpert 2008), which has a built‐in equipment economics
calculator. The calculator uses prices and correlations from the mentioned textbook by
Ulrich and Vasudevan. These prices were then scaled up using the appropriate cost
index.
For an economic analysis, a comparison between our design and an alternative was
completed. The rationale behind this type of analysis was that we were not able to do a
discounted cash flow rate of return on the design proposed since it is a sub‐system of a
larger process. Therefore, an economic feasibility study could not be done since all the
parts of the larger process are inter‐related and required to determine the economic
viability. As a result, a comparison was completed on the cost of energy required to
44
break the azeotrope of ethanol at 95%v/v in a distillation column versus using PSA.
6.2 Equipment Costs
As mentioned earlier, the costs for the equipment were calculated using methods
presented by Ulrich and Vasudevan in the program EconExpert. To scale up these
prices, a Chemical Engineering Plant Cost Index of 528.2, retrieved from the Chemical
Engineering Journals, was used. The scaled up costs of each individual piece of
equipment is summarized in Table 6.1. The total grass roosts capital was calculated to
be approximately $370,000. This cost breaks down into a total module cost and an
auxiliary facilities cost. Within the total module cost, it includes sub sections including
bare module costs as well as contingency and fee costs. The bare module costs take
into account freight, taxes, insurance, construction overhead, and engineering
expenses. The auxiliary facility costs include site development, auxiliary buildings, and
off‐site facilities. A break down of these sections and the respective costs are listed in
Table 6.1. The calculation of the contingency and fee costs is based on a percentage of
the total bare module cost where as the auxiliary costs are based on a percentage of the
total module cost. The respective percentages are listed in brackets behind the
applicable descriptions.
45
TABLE 6.1 Summary of equipment economics
Equipment Type Bare Module Cost
Heat Exchangers
U‐Tube $11,000
Bayonet (Heating) $33,000
Bayonet (Cooling) $27,800
Vacuum Pump Liquid Ring $49,500
Adsorption Tower
Tower 1 $59,500
Tower 2 $59,500
Total Bare Module Cost $240,000
Contingency(15%) and Fee(3%) $43,300
Total Module Cost $284,000
Auxiliary Facilities (30%) $85,100
Grass Roots Capital $369,000
46
6.3 Molecular sieve costs
The molecular sieve needs to be replaced approximately every 4000 cycles, which works
out to be about approximately every 3 months resulting in three bed changes per year.
The cost of the sieve is approximately $9 per kilogram. With each column having 600kg
in it, a cost per replacement is equal to $10,800 with a total cost of $32,400 per year.
6.4 Alternative Economic Comparison
To determine the cost of energy required in azeotropic distillation, the steam flow rate
within the system needed to be determined. A paper on saline extractive distillation
was first used to find the energy requirement to be approximately 5 MJ for breaking the
azeotrope. Comparing this value with the energy requirement for the proposed design,
it was found to have a ratio of 3 to 1. In other words, azeotropic distillation used three
times as much energy to break the azeotrope. Knowing this ratio, and that this design
requires a steam flow rate of approximately 4.2 , the steam flow rate required in
azeotropic distillation was found to be approximately 12.6 . Final costs were then
calculated using correlations out of Ulrich’s textbook again and determined to be
approximately $2.2 MM for PSA and $6.6 MM for azeotropic distillation. A summary of
the values needed in the calculation of the final costs can be seen in Table 6.2.
47
Table 6.2 Summary of alternative comparison
Pressure Swing
Adsorption Azeotropic Distillation
Aux. Boiler Steam Capacity, �s, (kg/s) 40 40
Pressure, P, (kPa) 250 250
Price of Fuel, CS,f, ($/GJ) 4.75 4.75
Plant Cost Index 528.2 528.2
Utility Price ($/kg) 0.0173 0.0173
Operating Time Per Year (s) 3.02E+07 3.02E+07
Steam Flow Rate (kg/s) 4.2 12.6
Annual Energy Cost $2,200,000 $6,610,000
*Sample calculations can be seen in Appendix E
48
Chapter 7.0: Safety Analysis
7.1 Introduction
Another important aspect of any design project is the evaluation of its safety. For this
project, it was specified that a safety analysis, including a hazard and operability
(HAZOP) analysis and a DOW Fire and Explosion Index analysis be performed on one
piece of equipment. The safety analysis was performed on one of the columns as it has
the most potential for problematic occurrences. The chemical properties of the system
were also analyzed by obtaining various material and safety data sheets (MSDS). In
order for this design to be deemed acceptable, every possible deviation that could
occur needed to be investigated and the proper preventative measures taken. This was
completed in the HAZOP analysis. A process safety management plan has also been
completed in efforts to make safety an important consideration to employees from the
beginning of the process and to continue to stress safety throughout its lifespan. Safety
is a very important part of any design and should not be taken lightly.
49
7.2 Chemical Properties
7.2.1 Ethanol
MSDS sheets were found for solutions of 95% v/v ethanol, 5% v/v water and for 100%
ethanol as these are the minimum and maximum concentrations of ethanol found in
our system. From these MSDS sheets it was found that the values for both
concentrations are equivalent. The threshold limit value of pure ethanol was found to
be 1000ppm. Although ethanol is a non‐reactive substance, it is very explosive and
flammable with the lower and upper flammable limits for both solutions being 3.3 %
and 19% respectively. It also has a very low flash point of 17.78 ˚C or 64 ˚F. On the
basis of a four hour exposure time, the acute oral toxicity (LD50) of ethanol was found
to be 3450kgmg
and the acute toxicity of the vapour (LC50) of ethanol was found to be
39000 3mmg
. All of these values can be found in Appendix G. This classifies ethanol to be
slightly hazardous in the case of ingestion or inhalation. It is also an irritant in the case
of skin contact.
Personal protective equipment such as protective gloves, safety goggles, and fire
retardant clothing are recommended whenever working with or near ethanol.
Concentrated ethanol solutions should be stored in a tightly closed and sealed
container in a cool well ventilated area that will not exceed a temperature of 23 ˚C.
Proper ventilation and accessible eye wash stations and dry chemical fire extinguishers
are also necessary.
50
7.2.2 Alumino Sillicate
An MSDS sheet was also found for the type 3A molecular sieve, which is alumino silicate
and can be found in Appendix G. This is a very stable, non‐flammable substance and is
only slightly hazardous in the case of inhalation. It is non‐toxic to humans with
exception to the case of chronic exposure which can possibly result in damage to the
lungs due to inhalation of dust formed by the molecular sieve. It can also be classified
as a slight irritant to the skin as it can react with moisture to create heat. This is not a
serious effect and can be easily removed with soap and water. The personal protective
equipment that is recommended when directly handling alumino silicate are safety
goggles, safety gloves, lab coat, and a dust respirator.
7.3 Hazard and Operability Analysis
7.3.1 HAZOP Strategy
A hazard and operability (HAZOP) analysis was completed on one of the towers. In this
analysis, all potential deviations that could occur on the tower were considered, rated,
and possible safeguards were implemented or recommended. This was done following
the method given in Ulrich’s Chemical Engineering: Process Design and Economics, A
Practical Guide. In this method, deviations were first identified and if possible
eliminated. If the deviation could not be eliminated it was minimized and isolated.
51
The HazardReview Leader2008 software, created by ABS consulting, was used to
conduct the identification portion of the HAZOP analysis. This software gives a list of
possible deviations for specific pieces of equipment. For the tower in this design, it
gave the following possible deviations; high temperature, low temperature, high
pressure, low pressure and leaks (ABS Consulting 2008). The possible causes and
possible consequences were then listed and analyzed and each deviation was then
rated, based on the severity of the consequences and the likelihood of the deviation
occurring. With this rating system, each deviation was given a value from 1 to 4 based
on the severity of its consequences, with 1 being the least severe, causing a single first
aid injury, and 4 being the most severe, causing multiple severe injuries. They were
then given a value from 1 to 4 based on the likelihood of occurrences, with 1 signifying
that it is not expected to occur ever, and 4 representing that it could potentially occur at
least once a year. Multiplying both values together gave the overall rating for the
deviation. Once the deviations were identified and rated, the proper precautionary
actions were taken to prevent and minimize the effects and likelihood of them
occurring.
7.3.2 HAZOP Conclusions
Using the strategy described above, each deviation was given a rating and analyzed.
This is summarized in Figure H.1. Here it can be seen that the most hazardous
deviations that could occur were determined to be high temperature on the tower and
52
a leak, with ratings of 4 and 8, respectively. These two should therefore be given
priority for preventative action.
High temperature on the column could be caused by either of the two proceeding heat
exchangers not working properly. If this were to occur, the pressure inside the column
would be increased, increasing the chances of condensation occur inside the column
and damaging the molecular sieve. Damage to the molecular sieve has no safety
related consequences itself, however, replacing the molecular sieve can be a potentially
hazardous job as it requires direct exposure to the ethanol and the molecular sieve. If
the temperature were to reach the auto‐combustion temperature of ethanol (363˚C) an
explosion could occur. Even though this is not expected to ever occur in the lifetime of
the system, the extreme nature of the consequences makes this a very hazardous
deviation. In order to prevent this, a high temperature alarm that will sound at 200˚C
has been included on the column. Temperature controls have also been put on the
second of the proceeding heat exchangers. These controls can be seen on Figure F.1.
A leak could possibly occur due to corrosion, rundown of equipment or wear on the
seals and connections to the tower. If this were to happen, the gaseous ethanol could
escape into the atmosphere and there were be a risk of explosion if it were to come into
contact with an ignition source. In order to prevent and detect a leak, leak detection
monitors, such as LEL monitors, and low pressure alarms and controls have been
implemented. These can be seen on Figure F.1. The pressure controls are also useful to
53
control the vacuum pump to make certain the regeneration column is being
depressurized adequately.
As seen in Figure F.1 flow rate transmitters have been placed on streams, 12, 18, 20 and
29 to control splitters 5 and 6. These controls were added to ensure that 40 % of the
product be sent to the regenerating tower to aid in the regeneration process. These
splitters are also controlled by a set of controls that measures the concentration of
water in stream 25 leaving the regeneration tower. Once the concentration in this
stream reaches zero, it is no longer necessary to purge the desorption tower with the
product. At this point the splitters are changed to ensure that 100 % of the product is
sent to the final heat exchanger for cooling.
The cooling water being fed into the third heat exchanger is regulated by temperature
controls that are signalled by a temperature transmitter on the final product stream 17.
This is to make sure that the product is cooled enough to completely condense it to a
liquid form. One of the stipulations of this design was that the liquid product stream
was to have a flow rate of 1250 . To satisfy this condition, flow transmitters have been
implemented on the outlet stream 14 that will signal valve 1 to open or close
accordingly.
54
In Figure F.1, it can also be seen that valves have been placed in the system so as to
make the isolation of every piece of equipment possible. This is necessary to carry out
maintenance, replacement of equipment, and for in the case of emergencies.
7.4 DOW Fire and Explosion Index Analysis
A Fire and Explosion Index analysis was completed on one of the towers. This
concluded that this adsorption system had a fire and explosion index of 95.5 which
means that it is a moderate hazard. The radius of exposure was found to be
approximately 80ft. The total value of equipment for this system is about $400,000,
however, because the value of the rest of the facility is unknown, an extra $10,000,000
has been added to the value of area of exposure to account fort the rest of the facility.
This gives a base and actual maximum property damage (MPPD) of $6,000,000 and
$4,360,000 respectively. The MPDO and Business Interruption Loss were found to be
15 days and $230,000 respectively.
7.5 Process Safety Management
As part of a process safety management plan, a monthly safety meeting will be held
that is compulsory for all employees to attend. These meetings will consist of the
discussion of any incidents or near‐misses that may have occurred in the previous
month. There will be a monthly safety theme that will be researched and presented by
the chair of the meeting. Each meeting will also have an allotted time slot for anyone
55
present to bring forward any safety concerns they may have. Scheduled preventative
maintenance and scheduled shutdowns will also be carried out as seen necessary.
The training for new employees will have a very safety oriented outlook, with job
shadowing, operating manuals and videos, and the completion of certain safety
courses. It is recommended that all employees complete the WHMIS, basic fire
extinguishing, confined space, and first aid courses before starting work with this
process.
56
Chapter 8.0: Conclusions
After conducting extensive research, performing calculations, and testing simulations,
Halo Consulting has determined that pressure swing adsorption is the best solution for
dehydrating a feed of 95%v/v ethanol to a final product of 99.5%v/v ethanol. The
design would consist of three heat exchanges, one liquid ring vacuum pump, and two
identical adsorption towers filled with a Type 3A molecular sieve.
The towers were found to be more efficient when tall and slender. Thus, the
dimensions of the towers are 4.5 m in height with a bed diameter of 0.55m and a wall
thickness of 4 mm. The height of the adsorption bed was found to be 3.5 m. The total
cycle time for one tower was approximately 48 minutes, 24 minutes for adsorption, 6
minutes for regeneration and 18 minutes for stand‐by, pressurization, and blow down.
During the 6 minutes of regeneration, 40 % of the dry ethanol gas product is sent to the
regenerating column to aid in desorption.
57
The areas of the three heat exchangers used in this process were calculated to be 1.96,
15.2, and 10.0 m2. The vacuum pump was sized to be 21.7 kW.
After an economic comparison, it was established that the annual cost of energy
required for azeotropic distillation was approximately three times that of pressure
swing adsorption, being approximately $6.6 MM and $2.2 MM respectively. The total
module cost and the overall grass roots capital for the pressure swing adsorption design
were calculated to be approximately $284,000 and $369,000, respectively.
Finally, after completing an analysis for both HAZOP and Dow’s Fire and Explosion Index
on one of the towers, it was determined that the main deviations of the system were
leaks and high temperatures in the column and that the system was classified as a
moderate hazard. Due to the flammable and explosive nature of ethanol, a process
safety management plan has also been implemented.
58
Chapter 9.0: Recommendations
After completing this design process, the following recommendations have been
suggested.
1. Gathering of experimental data at these conditions to observe the actual
decline in adsorption rates over time is strongly recommended. This data
would be used to perform a scale up operation of the design.
2. It is recommended that a simulation package with adsorption processes to
simulate the pressure swing adsorption. This could not be done for this
design as time did permit learning a new simulation program.
59
References
Aliasso, Joe. How to Size Liquid Ring Vacuum Pumps. Pumps and Systems Magazine.
2003, 1‐3. <http://www.graham-mfg.com/downloads/212.pdf>
Africa, Michael; Kendrick, Robert; Scramlin, Jeff; Catalano, Sam; Messacar, Julie; Palazzolo, Joseph.
Chemical Engineering Equipment. Macromedia, Inc. 1996.
Basmadjian, Diran. The Little Adsorption Book: A Practical Guide for Engineers &
Scientists. CRC Press, Inc: 1997.
Change, Hua; Yuan, Xi‐Gang; Tian, Hua; Zeng, Ai‐Wu. Experiment and prediction of
breakthrough curves for packed bed adsorption of water vapour on cornmeal.
Elsevier B.V. 2006, 1‐8.
Dickenson, T. Christopher. Valves, Piping and Pipelines Handbook. 3rd Ed. Elsevier
Advanced Technology. 1999.
EconExpert. Ulrich, Gael D.; Vasudevan, Palligarnai. February 2008.
< www.ulrichvasudesign.com>.
Graco Homepage. Graco Liquid Control. January 2008.
<http://www.graco.com/LCC/etoolbox/vacuum.html>.
60
HazardReview Leader 2006 Software. ABS Consulting. January 2008 – April 2008.
<http:/www.absconsulting.com/TC/103.html>.
Henley, Ernest J.; Seader, J. D. Separation Process Principles. 2nd Ed. . John Wiley &
Sons: New York, 2006.
Kirk, Othmer. Concise Encyclopedia of Chemical Technology. 4th Ed. John Wiley &
Sons: New York, 1999.
March Consulting Associates Inc. Homepage. March Consulting Associates Inc.
September 2007 – April 2008. http://www.marchconsulting.com.
Nevers. Noel De. Fluid Mechanics for Chemical Engineers. 3rd Ed. McGraw – Hill:
New York, 2005.
Perry, Robert H.; Green, Don W. Perry’s Chemical Engineer’s Handbook. 7th Ed.
McGraw – Hill: New York, 1997.
Peters, Max S. Elementary Chemical Engineering. 2nd Ed. McGraw – Hill Book
Company: New York, 1984.
Pinto, R.T.P.; Wolf‐Macial, M.R.; Lintomen, L. Saline extractive distillation process for
ethanol purification. Elsevier B.V. 2000, 1‐6.
61
Ruthven, Douglas M.; Farooq, Shamsuzzaman; Knaebel, Kent S. Pressure Swing
Adsorption. VCH Publishers, Inc, 1994.
Suzuki, Motoyuki. Adsorption Engineering. Kodansha Ltd, 1990.
Tien, Chi. Adsorption Calculations and Modeling. Butterworth‐Heinemann, 1994.
Ulrich, Gael D.; Vasudevan, Palligarnai. Chemical Engineering: Process Design and
Economics: A Practical Guide. 2nd Ed. Process Publishing: Durham, New
Hampshire, 2004.
ZEOCHEM Homepage. Zeochem. September 2007 – April 2008.
<http:/www.zeochem.com>.
62
Appendix A:
Process Flow Diagrams
63
Figure A .1: Process Flow Diagram mimicking the dehydration system in HYSYS when bed 2 (BAL‐2) is in regeneration
64
Figure A .2: Process flow diagram mimicking the dehydration system in HYSYS when bed 2 is done regenerating
65
Figure A.3 Process flow diagram of the ethanol dehydration system
66
Appendix B: Mass Balances
67
Table B.1 Mass balance for adsorbing column (Bed 1) time span for balance 0.106 h
Streams Mole Fraction Flowrate (kgmole/h)
Amount (kgmole)
Water Ethanol Feed Water Ethanol Feed Water EthanolIN 0.14 0.86 41.86 6.04 35.81 4.42 0.64 3.78 OUT 0.02 0.98 36.39 0.58 35.81 3.84 0.06 3.78 ACCUMULATED (adsorption rate) 1 0 5.47 5.47 0.00 0.58 0.58 0.00
time span for balance 0.293 h
Streams Mole Fraction Flowrate (kgmole/h)
Amount (kgmole)
Water Ethanol Product Water Ethanol Feed Water EthanolIN 0.14 0.86 25.11 3.63 21.48 7.37 1.06 6.30 OUT 0.02 0.98 21.84 0.35 21.48 6.41 0.10 6.30 ACCUMULATED (adsorption rate) 1.00 0.00 3.27 3.28 0.00 0.96 0.96 0.00
time span for balance 0.399 h = 23.94 Adsorption Time
Stream Amount (kgmole)
Total Water Ethanol IN 11.79 1.70 10.08 OUT 10.25 0.16 10.08 ACCUMULATED (adsorption rate) 1.54 1.54 0.00
68
Table B.2 Mass balance for desorbing column (Bed 2)
time 0.11 h
Streams Mole Fraction Flowrate (kgmole/h)
Amount (kgmole)
Water Ethanol Feed Water Ethanol Feed Water EthanolIN 0.02 0.98 14.56 0.23 14.33 1.54 0.02 1.51 GENERATED (desorption rate) 1.00 0.00 14.58 14.58 0.00 1.54 1.54 0.00 OUT 0.51 0.49 29.14 14.81 14.33 3.08 1.56 1.51
Stream Amount (kgmole)
Total Water Ethanol IN 1.5366284 0.024366 1.512262083OUT 3.076000272 1.563738 1.512262083GENERATED (desorption rate) 1.54 1.54 0.00
69
Table B.3 Mass leaving system
Time (hours)
Mole Fraction Flowrate (kgmole/h)
Amount (kgmole)
Water Ethanol Product Water Ethanol Product Water Ethanol 0.11 0.02 0.98 21.84 0.35 21.49 2.31 0.04 2.27 0.29 0.02 0.98 21.84 0.35 21.49 6.41 0.10 6.31
Total 8.71 0.14 8.58
Table B.4 Overall system mass balance
Stream Substance (kgmole)
Water Ethanol Total IN (feed) 1.70 10.08 11.79
‐ OUT (Product) 0.138 8.58 8.71 ‐ OUT (By‐Product) 1.56 1.51 3.08
= 0.00 0.00 0.00
70
Appendix C:
Adsorption Data
71
Figure C.1 Isothermal data for water adsorption on a type 3A molecular sieve
72
Figure C.2 Water vapour isotherm at 120˚C for type 3A molecular sieve
0
0.2
0.4
0.6
0.8
1
1.2
1.4
0 10 20 30 40 50 60
Water load
ing in % wt. of m
olecular sieve
Vapor Pressure in kPa
73
Figure C.3 Graph for the determination of equilibrium constant using Langmuir's form
y = 0.048x + 0.039R² = 0.998
0
0.5
1
1.5
2
2.5
3
0 10 20 30 40 50 60
ratio of partial pressure an
d water load
ing, p/q
Partial Pressure, p (kPa)
74
Table C.1 Table of given and calculated data for determining the breakthrough curve
Value Units Description MW mix, Mab 25.90 g/mol 3‐37 Henley Diffusion Volume ethanol, νa 51.17 Table 3.1 Henley Diffusion Volume water, νb 13.1 Table 3.1 Henley Diffusivity, Dab 1.03E‐01 cm2/s 3‐36 Henley 1.03E‐05 m2/s Effective Diffusivity, Deff 1.08E‐02 cm2/s 15‐75 Henley
1.08E‐03 m2/s
Average molecular velocity, νi 4.45E+02 14‐20 Henley Knudsen diffusion, Dk 4.75E‐10 cm2/s 14‐19 Henley 4.75E‐14 m2/s Surface Diffusion, Ds 9.40E‐05 cm2/s 15‐76 Henley 9.40E‐09 m2/s
Reynolds Number, NRe 7.21E+02 Table 3.3 Henley Schmidt Number, NSci 2.60E‐01 Table 3.3 Henley
Sherwood Number, Nsh 1.23E+01 15‐62 Henley Mass Transfer Coefficient, kc 4.20E‐02 m/s 15‐60 Henley 4.20E+00 cm/s
overall mass transfer coefficient, k 3.39E‐01 s‐1 15‐106 Henley
K 245 See Isotherm Sheet
75
Appendix D:
Sizing Calculations
76
Column Sizing:
1. Calculation of max velocity through column
.
.
. .
137 .
0.699
2. Calculation of the column diameter:
547.3 3600
0.6990.217
.
0.525 0.55
3. Calculation of the length of equilibrium zone:
77
. .
. .
1.92 2.00
Breakthrough Curve Calculation:
‐sub A is ethanol, sub B is water
1. Calculation of the molecular weight of the mixture:
21 1
21
46.07 /1
18.01 /25.9
2. Determination of the Diffusion volume of ethanol:
2 6 1
2 15.9 6 2.21 1 6.11
51.17
3. Calculation of the Diffusivity of ethanol in water:
0.00143 .
. ∑ ∑
0.00143 393.15 .
2.47 25.9 .51.17 13.1
78
0.0888
4. Calculation of the average molecular velocity:
8 / 8 8.314 / · 393.1542.02 /
/
445.1
5. Calculation of the Knudsen Diffusion Coefficient:
33.2 10 445.1 /
3 4.75 10
6. Calculation of the Surface Diffusivity:
1.6 100.45
1.6 100.45 990
18.01
2 1.987· 393.15
9.40 10
7. Calculation of the Reynolds Number:
0.003 2.057 ·
8.56 10 ·
7.21 10
79
8. Calculation of the Schmidt Number:
8.56 10 ·3.21 0.0888
0.3
9. Calculation of the Sherwood Number:
2 0.60 2 0.60 7.21 10 0.3 1.28
10. Calculation of the Mass Transfer Coefficient:
1.28 0.0888
0.0033.78
11. Determination of breakthrough time:
1
0.339 245 2
1.28
1 0.30.3
302
0.339 12902.0
1.28
2.50
1 erf √√
80
1 erf √2.50 √302√ . √
0.999
Solver for a c/cF = 0.068 for a 2 meter bed gave adsorption time of around 13.15 minutes. Length scaled up to 3.5 meters which gave adsorption of around 23.9 minutes.
81
Equipment Sizing:
Heat Exchanger – First Bayonet (heating)
1. Calculation of sensible heat area:
1a. Calculation of heat capacity of mixed stream:
% %
4.174·
0.0619 2.46·
0.9381
2.57·
1b. Calculation of heat duty:
∆
1759 2.57·
96.6 85.0
52364.7
1c. Calculation of log mean temperature:
∆
457 369.9
87.4
∆
456.6 358.0
98.6
82
∆ ∆ ∆∆∆
. .
.
.
92.89
1d. Calculation of sensible heat area:
∆
52365
1 900 · · 92.89
0.174
2. Calculation of the phase change area:
2a. Calculation of specific heat mixture:
% %
2260 0.0619 855 0.9381
941.97
83
2b. Calculation of heat duty:
1759 941.97
1657
2c. Calculation of log mean temperature:
∆
457 393
64
∆
456.6 358.0
98.6
∆ ∆ ∆∆∆
.
.
80.1
84
2d. Calculation of phase change area:
∆
1656924
1 900 · ·3600
1000 80.1
14.61
3. Calculation of superheating area:
3a. Calculation of heat capacity of mixed stream:
% %
4.174 · 0.0619 2.46 · 0.9381
2.57 ·
3b. Calculation of heat duty:
∆
1759 2.57·
393 369.6
105632.3
85
3c. Calculation of log mean temperature:
∆
457 393
64
∆
456.6 369.6
87
∆ ∆ ∆∆∆
74.91
3d. Calculation of superheating area:
∆
105632.3
1 900 · ·3600
1000 74.91
0.435
86
4. Calculation of total heat exchanger area:
0.174 14.61 0.435
15.22
Vacuum Pump
1. Calculation of the pump capacity:
15.92 3.2808 2
10 60 60
324.72
2. Calculation of the shaft work:
1
1
0.257 393.15 1.139 101.3
50
..
1
38.09 1.139 1 0.75
21.73
87
Table D.1 U‐Tube heat exchanger calculated data for pre‐heating feed
Value Units Cp,mix 2.57 kJ/kmol*KDelta T 50.00 K Heat Duty (Q) 225709.97 kJ/h ΔT1 35.00 K ΔT2 53.21 K ΔTLM 43.47 K A= 2.21 m2
Tables D.2. Bayonet heat exchanger calculated data for feed vaporization
Table D.2a Calculated data for sensible heating
Value Units Cp,mix 2.57 kJ/kg*K Delta T 11.60 K Heat Duty (Q) 52364.71 kJ/h ΔT1 87.40 K ΔT2 98.60 K ΔTLM 92.89 K A= 0.17 m2
Table D.2b Calculated data for phase change
Value Units λ,mix 941.97 kJ/kmol Delta T 35.00 K Heat Duty (Q) 1656924.35 kJ/h ΔT1 64.00 K ΔT2 98.60 K ΔTLM 80.06 K A= 14.61 m2
88
Table D.2c Calculated data for superheating
Value Units Cp,mix 2.57 kJ/kmol*KDelta T 23.40 K Heat Duty (Q) 105632.27 kJ/h ΔT1 64.00 K ΔT2 87.00 K ΔTLM 74.91 K A= 0.44 m2
Tables D.3 Bayonet heat exchanger calculated data for product condensation
Table D.3a Calculated Data for phase change
Value Units λ,mix -941.97 kJ/kmol
Heat Duty (Q) -
938390.02 kJ/h ΔT1 65.89 K ΔT2 20.00 K ΔTLM 38.49 K A= 8.69 m2
Table D.3b Calculated Data for sensible cooling
Value Units Cp,mix 2.47 Delta T -58.21 K
Heat Duty (Q) -
143280.09 kJ/h ΔT1 20.00 K ΔT2 65.89 K ΔTLM -38.49 K A= 1.33 m2
89
Table D.4 Calculated Data for liquid ring vacuum pump
Value Units Size 191.116 ft3/min 5.412 m3/min 324.72 m3/h Shaft Work(ws) 21.733754 kW Fluid Power 16.300315 kW
90
Appendix E:
Economics Calculations
91
Cost Summary The cost index is 528.2 Heat Exchangers : Shell and Tube : Fixed tube sheet and U-tube Total purchased cost = $ 3461 Material factor = 1.00 Pressure factor = 1.00 The bare module cost is = $ 11006 Heat Exchangers : Shell and Tube : Bayonet Total purchased cost = $ 10391 Material factor = 1.00 Pressure factor = 1.00 The bare module cost is = $ 33044 Heat Exchangers : Shell and Tube : Bayonet Total purchased cost = $ 8733 Material factor = 1.00 Pressure factor = 1.00 The bare module cost is = $ 27772 Gas Movers and Compressors : Blowers and compressors (cost of drive excluded) : Liquid ring Total purchased cost = $ 22517 The bare module cost is = $ 49538 Process Vessels (including towers) : Vertically oriented : With adsorbents, ion-exchange/chromatographic resins, catalysts Total purchased cost = $ 11878 Material factor = 1.00 Material factor for tower packing = 1.73 The bare module cost of tower packing is = $ 8552 The bare module cost is = $ 59459 Process Vessels (including towers) : Vertically oriented : With adsorbents, ion-exchange/chromatographic resins, catalysts Total purchased cost = $ 11878 Material factor = 1.00 Material factor for tower packing = 1.73 The bare module cost of tower packing is = $ 8552 The bare module cost is = $ 59459 *This was determined from econ expert and does not include replacement beds
92
Economics of an Azeotropic Distillation
1. Calculation of utility cost coefficient “a”:
2.3 10 .
2.3 10 40.
8.32 10 units?
2. Calculation of utility cost coefficient “b”:
3.4 10 .
3.4 10 2.5 .
3.56 10 ?
3. Calculation of utility price:
8.32 10 528.2 3.56 10 4.75$
0.01735$
4. Calculation of the operating time per year:
350 24 3600
93
3.02 10
5. Calculation of the annual cost:
0.01735$
3.02 10 12.6
6,610,000$
94
Appendix F:
Piping and Instrumentation Diagram
95
Figure F.1 (1) Temperature controls to control HX #2; (2)Pressure controls for leak detection and to ensure feed is at a high enough pressure to prevent condensation inside the column; (3) Pressure controls for leak detection and to control vacuum pump; (4) Concentration transmitters to regulate the use of the purge stream; (5) Flow rate controls to control the splitters 5 and 6; (6) Concentration transmitters to ensure product quality; (7) Flow rate controls to control output rate; (8) Temperature controls to regulate the amount of cooling water to HX #3.
96
Appendix G:
Material Safety Data Sheets
97
98
99
100
101
102
103
104
105
106
107
108
109
110
111
Appendix G:
Material Safety Data Sheets
T
Halo Consulting
Method: HAZOP
Team Members: Ma
No.: 2 Tow
Drawing: PID
Item Devia
2.1 High tempera
2.2 Low tempera
2.3 High pre
2.4 Low pres
Pl
Type: Column
ark Baier, Jamie Hiltz,
wer (D-120)
ation
ture HX #1 not
HX #2 not
ture HX #1 not
HX #2 not
essure High temp#1
High temp#2
High flow f
Block in prcolumn
ssure Low tempe#1
Low tempe#2
ant:
n
, Zack Taylor
Causes
t working properly
t working properly
t working properly
t working properly
perature from HX
perature from HX
from supply
rocess after
erature from HX
erature from HX
Site
Consequence
Increased pressure incolumn
High enough(500C) dto sieve
Auto-combustion of e= 363C
Possibility of condensin column
Damage to molecula
Possibility of condensin column
Reduced product
112
e:
Design Intent: Pa
es Cat
n
damage 4
ethanol 4
sation
r sieve
sation
Unit:
acked bed of aluminos
S L R
1 3 3 Te#2
Hig(at
4 1 4
4 1 4
2 1 2 Te#2
Loco
1 2 2 Pre(V-
Preco
Flo5)
Higstr
2 1 2
1 1 1 Pre(V-
Preco
Tower #1
silicate to be used to
Safeguards
emperature control on2
gh temp alarm on colt 200C)
emperature control on2
w temp alarm on thelumn (100C)
essure control on valv-2)
essure relief valve on lumn
ow control on splitter
gh pressure alarm onream entering column
essure control on valv-2)
essure relief valve on lumn
System
dehydrate Ethanol
n HX
umn
Rec 2. Considcontrol on HX
n HX
Rec 2. Considcontrol on HX
ve
(S-
n
ve
m: Dehydration of E
Action Items
der putting temperatuX #1
der putting temperatuX #1
thanol
ure
ure
No.: 2 Tow
Drawing: PID
Item Devia
2.5 Loss of containm(leak)
Figure H.1 Summ
wer (D-120)
ation
ment Rundown o
Corrosion
Wear on setower
mary of HAZOP an
Causes
of equipment
eal/connections to
nalysis
Consequence
Leakage of ethanol inatmosphere
Lower product
Potential explosion
Decreased pressure column; possible condensation
113
es Cat
nto
4
on
S L R
Lostr
Flo5)
2 2 4 Le(LE
1 2 4
4 2 8
1 2 2 Lostrco
Safeguards
w pressure alarms onreams entering colum
ow control on splitter
ak detection monitorsEL’s etc)
w pressure alarms onreams entering the lumn
n mn
(S-
s
n
Action Items
Top Related