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Transcript of Development of a Condenser for Marine Florae Pyrolysis Reactor
Development of a Condenser for the Developed
Marine Florae Pyrolysis Reactor
A masteral thesis presented to the
Department of Mechanical Engineering
University of San Carlos
In Partial Fulfillment of the Requirements
For the Degree of
Master of Engineering in Mechanical Engineering
By
Richard Jess L. Chan
October 2011
Edwin A. Carcasona, Ph.D.
Thesis Adviser
i
Acknowledgement
First and foremost, I would like to give thanks to the almighty God for the gift of
knowledge and strength, and all the blessings that He has showered to my life. I would
also like to express my gratitude the following for their contributions to this thesis:
To the Department of Science and Technology (DOST) and the Engineering Research &
Development for Technology (ERDT) program for funding my masteral education and
this research.
To my thesis adviser Dr. Edwin Carcasona for his much needed guidance and knowledge
regarding the research topic.
To my panelists Dr. Nicanor Buenconsejo, Dr. Ronald Galindo, and Engr. Joey Pastoril
whose comments and criticisms have led to the enrichment of this paper.
To the proponents of the other theses conducted in parallel with this thesis: Felixberto
Esgana Jr., Ivan Jhove Pacul, Michelle Rose Signe, Julius Enrico Valencia, Kenny
Alberto, Vhon Alfer Alivio, Junald Lasquites, and Eduard Tangub. Your help during the
wearisome experimentation is sincerely appreciated. And to John Paul and Manong Eddie
who also contributed to the completion of this thesis.
To my entire family; for their love and support that helped me to finish this thesis. And to
Mae Allequir who was the source of my inspiration and vigor during the making of this
paper.
Lastly, to all my friends and acquaintances that I have failed to mention but in some way
had contributed to the accomplishment of this thesis. Thank you very much!
ii
Development of a Condenser for the Developed Marine Florae Pyrolysis Reactor
By: Richard Jess L. Chan
Abstract
Two double-pipe condensers were designed, fabricated and tested for separating
the bio-oil from the pyrolysis gas of marine florae after pyrolysis reaction has occurred.
The inner tube of one condenser was made from Aluminum and the other from Stainless
Steel, both having a nominal diameter of 1 in. Initially, the marine florae pyrolysis
products were assumed to be similar to that of other biomass. These assumptions were
used to calculate the initial required lengths of the condensers. The calculations yielded
lengths of 78.1 cm and 99.9 cm for the aluminum and stainless condensers, respectively.
The fabricated condensers were tested by connecting it to the developed marine florae
pyrolysis reactor and conducting an actual pyrolysis experiment. The pyrolysis products
yield was determined. The maximum rate of bio-oil yield was found to be 5.33 ml/min
and the pyrolysis gas components that were determined were CO2 and CH4. The bio-oil
was also observed to stick to the walls of the condenser. Among the two condensers, the
aluminum condenser was easier to clean and had less oil that stick to its walls. The
required condenser length was recalculated by incorporating the determined rate of bio-
oil yield and pyrolysis gas components to the calculations. Also, the two-phase flow of
the volatiles was considered in the recalculation. The recalculated length was found to be
impractically too long for a double-pipe condenser. Other types of condensers, e.g. shell-
and-tube, are suggested for future studies.
Engr. Edwin A. Carcasona, PhD, PME
Thesis Adviser
iii
TABLE OF CONTENTS
Acknowledgement .............................................................................................................. i
Abstract .............................................................................................................................. ii
List of Tables ................................................................................................................... vii
List of Figures ................................................................................................................... ix
Chapter 1. Problem Setting and Background .................................................................1
1.1. Introduction .............................................................................................................1
1.2. Statement of the Problem ........................................................................................2
1.3. Significance of the Study ........................................................................................3
1.4. Objectives ...............................................................................................................4
1.5. Scope and Limitation ..............................................................................................4
1.5.1. Condenser Design ..........................................................................................5
1.5.2. Sources and Types of Marine Florae Feedstock ............................................6
1.6. Theoretical Background ..........................................................................................6
1.6.1. Pyrolysis of Marine Florae ............................................................................6
1.6.2. Condensation Phenomenon............................................................................8
1.6.3. Flow Rates .....................................................................................................8
1.6.4. Conservation Laws ........................................................................................9
1.6.5. Heat Transfer ...............................................................................................10
1.6.6. Gas Mixtures ................................................................................................13
1.6.7. Dimensionless Numbers ..............................................................................14
1.6.8. Homogeneous Two-Phase Model ................................................................15
Chapter 2. Review of Related Literature.......................................................................18
2.1. Pyrolysis Of Biomass............................................................................................18
2.1.1. Bio-oil ..........................................................................................................19
2.1.2. Liquid Collection .........................................................................................20
2.1.3. Pyrolysis Gas ...............................................................................................20
2.2. Double-Pipe versus Other Types of Condensers ..................................................21
2.2.1. Shell-and-Tube.............................................................................................21
iv
2.2.2. Spiral-Tube ..................................................................................................22
2.2.3. Plate-Fin .......................................................................................................22
2.2.4. Gasketed Plate..............................................................................................23
2.2.5. Spiral Plate ...................................................................................................24
2.2.6. Direct Contact ..............................................................................................24
2.3. Condensers Used in Pyrolysis...............................................................................25
2.3.1. Unapumnuk (1999) .....................................................................................25
2.3.2. Mudolodu (2002) ........................................................................................25
2.3.3. Jih (1982) ....................................................................................................26
2.3.4. Añora (2010) ...............................................................................................26
2.4. Condensation of Mixtures .....................................................................................27
2.5. Research Gap ........................................................................................................28
Chapter 3. Methodology ..................................................................................................29
3.1. Introduction ...........................................................................................................29
3.2. Condenser Design Process ....................................................................................30
3.2.1. Required Heat Transfer ................................................................................31
3.2.2. Convection Heat Transfer Coefficient .........................................................34
3.2.3. Logarithmic Mean Temperature Difference ................................................36
3.2.4. Heat Transfer Area.......................................................................................36
3.3. Marine Florae Collection and Preparation ............................................................37
3.4. Installation of Centrifugal Blower ........................................................................37
3.5. Experiment Set-up and Procedure ........................................................................39
3.5.1. Equipment Preparation ................................................................................41
3.5.2. Cooling Water Flow Calibration..................................................................42
3.5.3. Fluid Temperature Measurement .................................................................43
3.5.4. Periodic Oil Collection and Measurement ...................................................44
3.5.5. Static Pressure Measurement .......................................................................45
3.5.6. Gas Velocity Measurement ..........................................................................46
3.5.7. Gas Collection for Gas Chromatography.....................................................48
3.6. Condenser Evaluation ...........................................................................................48
3.6.1. Cleanability ..................................................................................................48
v
3.6.2. Pressure Drop ...............................................................................................49
3.6.3. Actual Heat Transferred...............................................................................53
3.7. Recalculation of the Double-Pipe Condenser Length...........................................54
3.7.1. Properties of Bio-oil and Pyrolysis Gas .......................................................55
3.7.2. Mass Flux .....................................................................................................57
3.7.3. Required Heat Transfer ................................................................................58
3.7.4. Logarithmic Mean Temperature Difference ................................................59
3.7.5. Convection Heat Transfer Coefficients .......................................................60
3.7.6. Length of the Condenser ..............................................................................60
3.7.7. Pressure Drop ...............................................................................................61
Chapter 4. Results and Discussion .................................................................................62
4.1. Designed and Fabricated Double-Pipe Condenser ...............................................62
4.2. Temperature of Condenser Fluids.........................................................................63
4.2.1. Temperature of Volatiles .............................................................................63
4.2.2. Temperature of Cooling water .....................................................................66
4.3. Static Pressure and Gas Velocity ..........................................................................69
4.4. Bio-oil Yield .........................................................................................................71
4.4.1. Effect of Blower on Bio-oil Yield ...............................................................72
4.4.2. Bio-oil Leakage............................................................................................74
4.4.3. Black Viscous Liquid...................................................................................74
4.5. Pyrolysis Gas ........................................................................................................76
4.5.1. Components .................................................................................................76
4.5.2. Estimate of Pyrolysis Gas Yield ..................................................................77
4.6. Condenser Performance ........................................................................................77
4.6.1. Condenser Material ......................................................................................78
4.6.2. Pressure Drop ...............................................................................................79
4.6.3. Actual Heat Transferred...............................................................................79
4.7. Results of Recalculation of The Condenser Length .............................................80
4.7.1. Comparison of Initial Calculation and Recalculation ..................................80
4.7.2. Effect of Flow Velocity ...............................................................................81
4.7.3. Effect of Thermal Conductivity of Condenser Tube ...................................81
vi
4.7.4. Effect of Cooling Water ...............................................................................82
Chapter 5. Conclusions and Recommendations............................................................84
5.1. Conclusions ...........................................................................................................84
5.2. Recommendations .................................................................................................85
Appendices ........................................................................................................................88
Appendix A. Calculation of Initial Condenser Design ................................................88
A.1. Required Heat Transfer ..................................................................................88
A.2. Logarithmic Mean Temperature Difference ..................................................89
A.3. Convection Heat Transfer Coefficients..........................................................90
A.4. Condenser Length ..........................................................................................92
Appendix B. Fabricated Condenser Parts and Assembly ............................................93
B.1. Dimensions and Parts .....................................................................................93
B.2. Condenser Accessories ...................................................................................94
B.3. Thermocouple Probes and Pressure Taps.......................................................95
B.4. Condenser Tilt Angle .....................................................................................96
Appendix C. Cooling Water Flow Rate Measurements ..............................................97
Appendix D. Static Pressure and Gas Velocity Measurements ...................................98
D.1. Static Pressure Measurements ........................................................................98
D.2. Gas Velocity Measurements ..........................................................................99
Appendix E. Pyrolysis Products ................................................................................100
E.1. Periodic Bio-oil Volume Measurement ........................................................100
E.2. Bio-oil and Pyrolysis Gas Yield ...................................................................102
E.3. Product Composition and Residence Time from Añora (2010) ..................103
Appendix F. Volatile Temperature Graph .................................................................104
F.1. Volatile Temperature Graph with Plotted Periodic Bio-oil Yield ................104
F.2. Volatile Temperature Graph without Plotted Periodic Bio-oil Yield ...........115
Appendix G. Calculation of Pressure Drop and Actual Heat Transfer ......................119
G.1. Pressure Drop ...............................................................................................119
G.2. Actual Heat Transfer ....................................................................................124
Appendix H. Recalculation of Double-Pipe Condenser Length ................................127
vii
H.1. Bio-oil and Pyrolysis Gas Properties ...........................................................127
H.2. Mass Flux .....................................................................................................129
H.3. Required Heat Transfer ................................................................................131
H.4. Logarithmic Mean Temperature Difference ................................................132
H.5. Convection Heat Transfer Coefficients........................................................133
H.6. Condenser Length ........................................................................................136
H.7. Pressure Drop ...............................................................................................137
Definition of Terms ........................................................................................................141
Bibliography ...................................................................................................................143
LIST OF TABLES
Table 1.1: Types of Marine Florae Feedstock ...............................................................6
Table 3.1: Experiment Runs ........................................................................................41
Table 3.2: Necessary Bio-oil Properties ......................................................................56
Table 3.3: Necessary Gas Properties ...........................................................................56
Table 4.1: Cooling Water Temperature Reading for Run A2 ......................................67
Table 4.2: Cooling Water Inlet and Volatile Exit Temperatures for Run S1 ..............68
Table 4.3: Inlet Static Pressure ....................................................................................69
Table 4.4: Gas Velocity in the Condenser Inner-Tube ................................................70
Table 4.5: Collected Bio-oil for Run A7 .....................................................................72
Table 4.6: Component Percentage of Pyrolysis Gas....................................................76
Table A.1: Values of Variable in Eq. (23) ..................................................................92
Table B.1: List of Parts ................................................................................................93
Table B.2: Inner-Tube Actual Dimensions ..................................................................94
Table B.3: Outer-Tube Actual Dimensions .................................................................94
Table C.1: Mass Flow Rate for Fully Open .................................................................97
Table C.2: Mass Flow Rate for One Valve-Turn .........................................................97
Table C.3: Mass Flow Rate for Two Valve-Turns ......................................................97
viii
Table C.4: Mass Flow Rate for Three Valve-Turns ....................................................97
Table C.5: Mass Flow Rate for Four Valve-Turns ......................................................97
Table D.1: Manometer Reading for Run A4, „full open‟ ............................................98
Table D.2: Manometer Reading for Run A4, „slightly close‟ .....................................98
Table D.3: Manometer Reading for Run A5, „full open‟ ............................................98
Table D.4: Manometer Reading for Run A5, „slightly close‟ .....................................98
Table D.5: Gas Exit Velocity .......................................................................................99
Table D.6: Velocity Inside Inner-Tube ........................................................................99
Table E.1: Bio-oil Volume Collected for Run A1 .....................................................100
Table E.2: Bio-oil Volume Collected for Run A2 .....................................................100
Table E.3: Bio-oil Volume Collected for Run A3 .....................................................100
Table E.4: Bio-oil Volume Collected for Run A4 .....................................................100
Table E.5: Bio-oil Volume Collected for Run A5 .....................................................100
Table E.6: Bio-oil Volume Collected for Run A6 .....................................................100
Table E.7: Bio-oil Volume Collected for Run A7 .....................................................101
Table E.8: Bio-oil Volume Collected for Run A8 .....................................................101
Table E.9: Bio-oil Volume Collected for Run S5 ......................................................101
Table E.10: Bio-oil Volume Collected for Run S6 ....................................................101
Table E.11: Bio-oil Volume Collected for Run S7 ....................................................101
Table E.12: Bio-oil Volume Collected for Run S8 ....................................................102
Table E.13: Mass of Marine Florae Feedstock and Pyrolysis Products ....................102
Table E.14: Mass Percentage of Pyrolysis Products ..................................................102
Table E.15: Density of Bio-oil ...................................................................................103
Table E.16: Mass Percentage and Residence Time for Green Algae ........................103
Table E.17: Mass Percentage and Residence Time for Red Algae............................103
Table E.18: Mass Percentage and Residence Time for Brown Algae .......................103
Table E.19: Mass Percentage and Residence Time for Seagrass...............................103
Table G.1: Absolute Viscosities of Pyrolysis Gas Components ................................120
Table G.2: Summary of Pressure Drop for Run A4, „full open‟ ................................123
Table G.3: Summary of Pressure Drop for Run A4, „slightly close‟ .........................123
Table G.4: Summary of Pressure Drop for Run A5, „full open‟ ................................124
ix
Table G.5: Summary of Pressure Drop for Run A5, „slightly close‟ .........................124
Table G.6: Constants for Eq. (G.19) .........................................................................124
Table G.7: Summary of Heat Transfer for Run A4, „full open‟ ................................125
Table G.8: Summary of Heat Transfer for Run A4, „slightly close‟ .........................125
Table G.9: Summary of Heat Transfer for Run A5, „full open‟ ................................126
Table G.10: Summary of Heat Transfer for Run A5, „slightly close‟ .......................126
Table H.1: Bio-oil Properties Applied in Desuperheating Zone................................127
Table H.2: Bio-oil Properties Applied in Condensing Zone......................................127
Table H.3: Bio-oil Properties Applied in Subcooling Zone ......................................127
Table H.4: Constants for Eq. (H.1) ...........................................................................127
Table H.5: Pyrolysis Gas Properties Applied in Desuperheating Zone .....................129
Table H.6: Pyrolysis Gas Properties Applied in Subcoolnig Zone ............................129
Table H.7: Values of Variables in Eq. (H.33) ...........................................................135
Table H.8: Summary of Required Condenser Length ...............................................139
Table H.9: Summary of Pressure Drop ......................................................................140
LIST OF FIGURES
Figure 1.1: Double-Pipe Heat Exchanger ......................................................................2
Figure 1.2: Average Mass Loss Curve with respect to Time (without binder) .............7
Figure 1.3: Average Mass Loss Curve with respect to Time (with binder)...................7
Figure 1.4: Diagram of Double-Pipe Counter-Flow Heat Exchanger .........................11
Figure 1.5: Temperature Profile for Counter-Flow Heat Exchanger ...........................12
Figure 2.1: Shell-and-Tube Condenser ........................................................................21
Figure 2.2: Plate-Fin Condenser ..................................................................................22
Figure 2.3: Gasketed Plate Heat Exchanger ................................................................23
Figure 2.4: Experiment Set-up of Mudulodu (2002) ..................................................25
Figure 2.5: Experiment Set-up of Jih (1982) ..............................................................26
Figure 2.6: Experiment Set-up of Añora (2010) .........................................................27
Figure 3.1: Study Flow ................................................................................................29
x
Figure 3.2: Condenser Design Flow Chart ..................................................................31
Figure 3.3: Gas Escaping through the Feed Port of the Reactor..................................38
Figure 3.4: RPM and Air Velocity Measurement ........................................................38
Figure 3.5: Retrofitted Centrifugal Blower..................................................................39
Figure 3.6: Schematic of Experiment Set-up ...............................................................39
Figure 3.7: Actual Experiment Set-up without Manometer ........................................40
Figure 3.8: Insulated Condenser ..................................................................................42
Figure 3.9: Installed Condenser ...................................................................................42
Figure 3.10: Thermocouple Datalogger .......................................................................43
Figure 3.11: Condenser with Thermocouple Probes ...................................................44
Figure 3.12: Bio-oil Collection and Storage ................................................................44
Figure 3.13: Static Pressure Measurement Set-up .......................................................45
Figure 3.14: Inclination Positioning Instruments ........................................................45
Figure 3.15: Gas-Exit-Valve Positions ........................................................................47
Figure 3.16: Uro-bag filled with Pyrolysis Gas ...........................................................48
Figure 3.17: Temperature Profile.................................................................................55
Figure 3.18: Conservation of Mass in the Condenser ..................................................58
Figure 4.1: Condenser Length .....................................................................................62
Figure 4.2: Condenser Tilt Angle ................................................................................62
Figure 4.3: Volatile Temperature Graph of Run A1 ....................................................63
Figure 4.4: Temperature Rise while Blower was Turned Off for Run A4 ..................64
Figure 4.5: Volatile Exit and Cooling Water Inlet Temperatures for Run S1 .............65
Figure 4.6: Cooling water Exit Temperature for Run S1 ............................................67
Figure 4.7: Fan-System Curve .....................................................................................70
Figure 4.8: Collected Bio-oil .......................................................................................71
Figure 4.9: Volatile Temperature Graph......................................................................73
Figure 4.10: Bio-oil Leakage Plotted in Volatile Temperature Graph ........................74
Figure 4.11: Unrecovered Black Viscous Liquid ........................................................74
Figure 4.12: Collected Black Viscous Liquid ..............................................................75
Figure 4.13: Black Viscous Liquid Residue ................................................................75
Figure 4.14: Flame from Pyrolysis Gas .......................................................................76
xi
Figure 4.15: Comparison of Stainless and Aluminum Condensers .............................78
Figure 4.16: Aluminum Condenser .............................................................................78
Figure 4.17: Flow Velocity, Condenser Length, Pressure Drop ..................................81
Figure 4.18: Cooling Water Convection Coefficient and Condenser Length..............82
Figure B.1: Aluminum Condenser ...............................................................................93
Figure B.2: Stainless Condenser ..................................................................................93
Figure B.3: Exploded View of the Condenser .............................................................94
Figure B.4: Adapter .....................................................................................................94
Figure B.5: Static Pressure Tap ...................................................................................95
Figure B.6: Position of Thermocouple Probes .............................................................95
Figure B.7: Position of Static Pressure Taps ...............................................................95
Figure B.8: Condenser Tilt Angle ................................................................................96
Figure F.1: Volatile Temperature Graph of Run A1 ..................................................104
Figure F.2: Volatile Temperature Graph of Run A2 ..................................................105
Figure F.3: Volatile Temperature Graph of Run A3 ..................................................106
Figure F.4: Volatile Temperature Graph of Run A4 ..................................................106
Figure F.5: Volatile Temperature Graph of Run A5 ..................................................107
Figure F.6: Volatile Temperature Graph of Run A6 ..................................................108
Figure F.7: Volatile Temperature Graph of Run A7 ..................................................109
Figure F.8: Volatile Temperature Graph of Run A8 ..................................................110
Figure F.9: Volatile Temperature Graph of Run S5 ..................................................111
Figure F.10: Volatile Temperature Graph of Run S6 ................................................112
Figure F.11: Volatile Temperature Graph of Run S7 ................................................113
Figure F.12: Volatile Temperature Graph of Run S8 ................................................114
Figure F.13: Volatile Temperature Graph of Run S1 ................................................115
Figure F.14: Volatile Temperature Graph of Run S2 ................................................116
Figure F.15: Volatile Temperature Graph of Run S3 ................................................117
Figure F.16: Volatile Temperature Graph of Run S4 ................................................118
1
CHAPTER 1
PROBLEM SETTING AND BACKGROUND
1.1. Introduction
Energy is a resource that the modern society cannot live without. However,
studies have shown that the conventional method of harnessing energy, e.g. burning of
fossil fuels, is literally steadily killing the environment. To prevent further destruction of
the environment attention has been given to renewable energy sources. One such
renewable energy resource, that is the topic of this study, is biomass, particularly marine
florae or more commonly known as seaweeds.
Earlier studies regarding the use of marine florae as an energy resource have been
conducted by Baring, et al (2009)[3]
and Añora (2010)[1]
and were able to present
promising results. Añora studied the extraction of useful fuel products from marine florae
by means of a method known as pyrolysis. This was replicated by Esgana (2011)[13]
on a
much larger scale. Pyrolysis is the heating of the biomass in the absence of oxygen to
produce solid, liquid, and gaseous end products such as carbonaceous char, bio-oil, and
pyrolysis gases, respectively.[25][26]
During pyrolysis reaction volatiles are released from
the biomass. These volatiles are composed of the condensable and noncondensable
components, which are the bio-oil and pyrolysis gas, respectively.
Esgana‟s study was to develop a marine florae pyrolysis reactor that was capable
of pyrolyzing much larger quantities of marine florae than in Añora‟s study. In any
pyrolysis system the pyrolysis reactor is coupled with a condenser for separating the
condensable from the noncondensable component of the volatiles, which is discussed in
references [1], [16], [20], [24]. The purpose of the present study was to develop the
condenser for the marine florae pyrolysis reactor. The pyrolysis experiment of the present
study was done simultaneously with Esgana‟s experiment. Since pyrolysis product yield
is also dependent on the heat rate and reactor conditions, the results of this study is
applicable only for the reactor developed by Esgana or other reactors of the similar
specifications. At present, there are no recognized design methods and most work has
been empirical and specific to the characteristics of the feedstock being processed.
2
Commercial liquids recovery processes are usually proprietary and may be specific to
individual feedstock, reactor configurations and products.[7]
Being the first condenser designed for the developed marine florae pyrolysis
reactor, the simplest type of condenser was seen as the best starting point. The type of
condenser chosen was the double-pipe heat exchanger, shown in Figure 1.1, because it is
easy to fabricate, maintain, and its tubular construction allows it to be easily “scaled up”
to shell-and-tube if greater heat exchange duties are required. Also, a tubular condenser is
effective in separating the oil from the pyrolysis gas.[9]
Cleanability of the condenser was
a major concern because the volatiles may contain solid char particles due to carry-over
from the pyrolysis reactor.[25]
The carry-over char was also observed in the present study.
It was also observed from the study of Añora (2010) that the bio-oil deposited to the
walls of the reactor which then required frequent cleaning. The performance of the initial
design was analyzed: flaws and problems were identified and improvements were
suggested.
Figure 1.1: Double-Pipe Heat Exchanger
1.2. Statement of the Problem
Design of condensers requires that the properties of both the cooling fluid and the
vapor to be condensed be known. However, at the beginning of this study the researcher
did not have data regarding the properties of the marine florae volatiles. Thus,
assumptions of the composition and properties of marine florae volatiles were required to
be able to come up with the initial condenser design. The assumptions were based on
literatures on the pyrolysis of other types of biomass (mostly wood).[8][25][26]
Properties of
the volatiles of other biomass could be far different from that of marine florae which
3
could result in the failure of the condenser operation: the calculated heat transfer area
might be insufficient, thus the required amount of heat rejection might not be attained
and the condensable component might not be condensed. Another problem that existed in
the design of the condenser was the presence of the noncondensable component.
Condensation of mixtures with noncondensable gases is a complicated phenomenon.[14]
Some assumptions were necessary in order to simplify the calculations but, in the
process, sacrificed the accuracy of the solution.
Another factor that was considered in condenser design was the material. In the
case of corrosive fluids, one might need to use expensive corrosion-resistant materials
such as stainless steel or even titanium.[10]
In this study two types of materials were used
to construct the condenser, and were visually inspected for any signs of corrosion. The
two types of material were also tested for cleanability since it was observed in the
experiment of Añora (2010) that the bio-oil sticks to the walls of the distilling flask,
which was used as the pyrolysis reactor, and the glass condenser. This could cause
fouling in the condenser which would decrease the effectiveness. After the experiment,
the researcher inspected which condenser material had less bio-oil that stick to it and
which was more easily cleaned.
1.3. Significance of the Study
Literatures support that the volatiles obtained from biomass are good candidates
as alternative fuels.[2][21][23]
Since the volatiles are composed of the condensable and
noncondensable components, a condenser is needed to condense the condensable
component and separate it from the noncondensable component. There are many
parameters that must be considered in condenser design. Some of these parameters are
properties of the marine florae volatiles, volatile flow rate, and condenser operating
conditions. Most of these parameters were unknown prior to the present study.
This study designed and fabricated a condenser based only on assumptions of the
design parameters. The purpose of the fabricated condenser was to expose an actual
condenser to the operating conditions encountered in the pyrolysis of marine florae using
the reactor developed by Esgana. In this manner some of the design parameters were
4
revealed during the experiment and after analysis of the acquired data. The present study
also investigated if the condenser material had any significant effect.
There is concrete theory on condenser design; however, good knowledge of the
environment in which the condenser will operate must be known to be able to achieve the
optimum design. Also, operating conditions vary with different systems in which the
condenser is used; hence, each condenser design is unique to its own system. In
pyrolysis, the amount of volatiles and volatile flow differ with type of feedstock and
reactor design. Since the reactor used in the present study was a new design developed by
Esgana[13]
, the actual operating conditions and the behavior of the volatiles is yet to be
known. Hence, investigation regarding the actual operating conditions was necessary.
Problems that arose in the experiment were identified and solutions were suggested. The
results of this study can be used as a bench mark for future condenser designs for the
developed marine florae pyrolysis reactor.[13]
1.4. Objectives
To design and fabricate two double-pipe condensers, one made from aluminum
and the other from stainless steel, for the marine florae pyrolysis reactor.
To evaluate the performances of the fabricated condensers.
To determine the percent amount of bio-oil and pyrolysis gas that can be extracted
and collected from a given marine florae feedstock.
To reconstruct the condenser design methodology based on the collected data
from the experiment.
1.5. Scope and Limitations
The scope of this study was to make a simple design, that is, the two-phase flow
of the volatiles was not considered, and to fabricate the condenser for the marine florae
pyrolysis reactor developed by Esgana. The condenser was used to separate the
condensable from the noncondensable component of the marine florae volatiles. The
researcher investigated how the condenser performed during the experiment. The amount
of volatiles extracted from the feedstock was also monitored, especially the bio-oil to aid
in the analysis of the condenser performance. Specifications/parameters for future
5
condenser design were derived based on the performance of the fabricated condenser and
other data obtained from the experiment.
This study was not concerned with the operation of the pyrolysis reactor.
However, data on the reactor temperatures were referred to whenever appropriate. Other
parts or components directly related to the operation of the reactor were not included in
this study.
1.5.1. Condenser Design
The condenser was fabricated using materials that were affordable and available
in the local market. Two double-pipe condensers were fabricated; one with an aluminum
inner tube and the other with stainless steel inner tube.
Very little was known about the marine florae volatiles, especially its
thermophysical properties which were necessary parameters in designing the condenser.
The properties of the volatiles were assumed based on literatures on other biomass
pyrolysis. The assumptions are discussed in Section 3.2 and in Appendix A together with
the details of the calculation.
The flow configuration that was tested in the experiment was counter flow only.
The volatiles flowed inside the inner tube and the cooling water in the annular space
between the inner and outer tubes. Counter flow was chosen over parallel flow because of
its superior heat transfer capability.[10]
The new design methodology was revised based on the obtained data from the
experiment. Only the condenser length was recalculated. The recalculation is discussed in
3.7 and Appendix H. This study was focused only on the thermal design of the condenser
and only little detail was given to the bio-oil collection system.
6
1.5.2. Sources and Types of Marine Florae Feedstock
The types of marine florae that were used as feedstock for the pyrolysis reactor
are Red, Brown, Green algae, and Seagrass only. Only drifted marine florae were used in
the experiment and collection was limited to the province of Cebu only. The types of
marine florae feedstock that were used in this study are listed in Table 1.1.
Table 1.1: Types of Marine Florae Feedstock
Green algae Brown algae Red algae Seagrass
Pelletized without
binder
Pelletized with
binder
Pelletized without
binder
Pelletized with
binder
Non-pelletized Pelletized without
binder Non-pelletized
Pelletized without
binder
1.6. Theoretical Background
The equations discussed in this Section are presented in their most basic form.
The equations take other forms through the discussions depending on the situation in
which they were used: type of fluid, flow condition (laminar of turbulent), geometry of
the flow area, etc. The applications of the equations presented here are discussed in
Chapters 3 and 4.
1.6.1. Pyrolysis of Marine Florae
The study of Añora (2010)[1]
had proven the possibility of extracting condensable
liquid (bio-oil) and combustible gas (pyrolysis gas) from marine florae. There were three
events that happened during pyrolysis experiments. First, from 0 minute to 4 minutes,
there was no change in the weight of the loaded sample pellets. Next event was the rapid
change of sample pellets‟ weight at 4 minutes to 35 minutes. A rapid bio-oil production
was observed from 4 minutes to 25 minutes. Also, start of pyrolysis gas production was
observed from 7 minutes to 15 minutes. Finally, a slow decrease of sample pellets‟
weight was observed from 35 minutes to 1 hour. It was suspected that pyrolysis reaction
has stopped since no production of bio-oil nor pyrolysis gas was observed.
7
Figure 1.2: Average Mass Loss Curve with respect to Time (without binder)[1]
Figure 1.3: Average Mass Loss Curve with respect to Time (with binder)[1]
Figure 1.2 and 1.3 shows the average mass loss curves with respect to time for
pellets with and without binder, respectively. The residence time of the experiment, from
start to the end of the experiment, are summarized in Appendix E. The percent product
compositions of bio-oil and pyrolysis gas are also shown in Appendix E. These data were
used to estimate for the mass flow rate of the volatiles which was then used to solve for
the required heat transfer in the initial design of the condenser.
8
1.6.2. Condensation Phenomenon
When condensing at low velocity, tube condensation is better in down-flowing
inclined tubes than either vertical or horizontal tubes. This is because the layer of
condensate in the bottom is quite thick in a horizontal tube, so that a small inclination in
the direction of flow results in more rapid condensate flow and much thinner condensate
layer. Vertical tubes are not usually as good as inclined ones because the condensate
layer is uniform around the tube; better heat transfer is obtained when the condensate
layer is nonuniformly distributed. The optimal inclination for condensation is about
20°.[14]
1.6.3. Flow Rates
Basic forms of the mass and volume flow rate[11]
are equations shown in Eq. (1.1)
and (1.2), respectively. These equations were mainly used to solve the flow rates of the
volatiles and the cooling water. Other forms of Eq. (1.1) and (1.2) were also used and
discussed in Chapters 3 and 4.
mm vA
t
1.1
mV vA
1.2
where: ṁ = mass flow rate, kg/s
V = volume flow rate, m3/s
ρ = density of fluid, kg/m3
v = velocity of flow, m/s
A = flow area, m2
Δm = net change in mass within a system, kg
Δt = elapsed time, sec
9
1.6.4. Conservation Laws
The conservation of mass principle[11]
, in Eq. (1.3), states that amount of mass
entering a system, such as a heat exchanger, minus the mass leaving it is equal to the
change of mass in the system.
in outm m m
1.3
where: Δm = net change in mass within a system, kg
min = total mass entering a system, kg
mout = total mass leaving a system, kg
Eq. (1.3) can also be expressed in the rate form as[11]
in out
dmm m
dt 1.4
where: dm/dt = rate of change of mass within a system, kg/s
min = total mass flow rate into a system, kg/s
mout = total mass flow rate out of a system, kg/s
For steady flow systems like heat exchangers, the rate of mass flow into the system must
be equal to the rate of flow out of it. Thus, Eq. (1.4) reduces to Eq. (1.5), which means
that the mass flow rate ṁ is constant.
in outm m m
1.5
10
The first law of thermodynamics states that, for steady state, steady flow systems,
namely heat exchangers, and neglecting heat losses, the heat released by the hot fluid is
equal to the heat absorbed by the cold fluid. In equation form,
released absorbedQ Q
1.6
where: Qreleased = heat released by the hot fluid, W
Qabsorbed = heat absorbed by the cold fluid, W
1.6.5. Heat Transfer
For steady-flow systems, such as heat exchangers, the rate of heat transfer is[10]
pQ m h mc T
1.7
where: Q = heat transfer, W
ṁ = mass flow rate, kg/s
Δh = change in enthalpy, J/kg
cp = specific heat at constant pressure, J/kg·K
ΔT = change in temperature, °C
The heat transfer in a heat exchanger is calculated from Eq. (1.8) [10]
lmTQ
R=
1.8
where: ΔTlm = logarithmic mean temperature difference, °C
R = total thermal resistance, °C/W
11
The total thermal resistance for a double-pipe heat exchanger is calculated from Eq. (1.9).
Figure 1.4 shows a diagram of a double-pipe heat exchanger. The LMTD is calculated
from Eq. (1.10).
ln1 1
2
o i
i i o o
d dR = + +
Ah πkL A h 1.9
where: Ai = inner surface area of the inner tube, m2
Ao = outer surface area of the inner tube, m2
di = inside diameter of the inner tube, m
do = outside diameter of the inner tube, m
L = length of the heat exchanger, m
k = thermal conductivity of heat exchanger material, W/m·K
hi = convection coefficient of fluid inside the inner tube, W/m2·K
ho = convection coefficient of fluid in the annular space, W/m2·K
Figure 1.4: Diagram of Double-Pipe Counter-Flow Heat Exchanger
12
The LMTD is
1 2
1 2
Δ Δ
ln Δ Δlm
T TT
T T
=
1.10
where ΔT1 and ΔT2 are illustrated in Figure 1.5.
Figure 1.5: Temperature Profile for Counter-Flow Heat Exchanger
The convection heat transfer coefficient discussed above can be determined from
Eq. (1.11) below.[10]
Nu
c
kh
L
1.11
where: Nu = Nusselt number, dimensionless
k = thermal conductivity of the fluid, W/m·K
Lc = characteristic length, m
The Nusselt number is discussed in Section 1.6.7. The characteristic length Lc is equal to
the tube diameter for flow inside tubes; in annular flow the hydraulic diameter is used in
place of Lc.
13
For film condensation of inside tubes, Eq. (1.12) is used to solve the convection
heat transfer coefficient. Eq. (1.12) is restricted to low vapor Reynolds number, Re <
35,000.
1/43
0.555
v fg
sat i
ρ ρ - ρ g'k h'h
μd T -T
1.12
where: 0.68fg fg w sat ih' h c T -T
sing' = g α
ρ = density of liquid film, kg/m3
ρv = density of vapor, kg/m3
μ = absolute viscosity of liquid film, kg/m·s
k = thermal conductivity of liquid film, W/m·K
hfg = latent heat of vaporization fluid, J/kg
c = specific heat of liquid film, J/kg·K
α = tilt angle of the condenser, deg.
d = inner diameter of the tube, m
Tsat = saturation temperature of vapor, °C
Ti = inner surface temperature of condenser wall, °C
g = acceleration due to gravity, 9.81 m/s2
1.6.6. Gas Mixtures
When a system is composed of more than one gas component, they can be
analyzed as a homogeneous mixture where its properties can be calculated from Eq.
(1.13).
n
mixture j j
j 1=
X y X
1.13
14
where: Xmixture = represents any property of the mixture
Xj = property of a single gas component j
yj = mass fraction of gas j
1.6.7. Dimensionless Numbers
The Reynolds number is used to characterize the flow regime in any fluid system.
The flow could either be laminar or turbulent, and is determined based on the value of
Reynolds number. In heat transfer calculations, specifically, different equations are used
to calculate the same parameter depending on the Reynolds number. The Reynolds
number can be calculated from Eq. (1.14).
Re cρvL
μ
1.14
where: Re = Reynolds number, dimensionless
ρ = density of the fluid, kg/m3
v = velocity of the flow, m/s
Lc = characteristic length of the flow channel, m
μ = absolute viscosity of the fluid, kg/m·s
In convection studies, it is common practice to nondimensionalize the convection
heat transfer coefficient with the Nusselt number defined as
Nu c
hL=
k
1.15
where: Nu = Nusselt number, dimensionless
h = convection heat transfer coefficient, W/m2·K
k = thermal conductivity of the fluid, W/m·K
Lc = characteristic length, m
15
1.6.8. Homogeneous Two-Phase Flow Model
The simplest approach to the treatment of the flow of a gas-liquid mixture in a
channel is to treat the flow as if the mixture were behaving as a homogeneous fluid, with
the velocities of the two phases identical. That is,
G L TPv v v
1.16
where: vG = velocity of the gas-phase, m/s
vL = velocity of the liquid-phase, m/s
vTP = velocity of the homogeneous mixture, m/s
With the assumption given above, the quality of the two-phase system is then given by
1
G G L
G G G L
x
1.17
where the void fraction εG is
GG
L G
V
V V
1.18
where: ρG = density of the gas-phase, kg/m3
ρL = density of the liquid -phase, kg/m3
GV = volume flow rate of the gas-phase, m3/s
LV = volume flow rate of the liquid-phase, m3/s
16
The density and absolute viscosity of the two-phase homogeneous mixture are given by
Eq. (1.19) and (1.20) below, respectively.
1
G LTP
L Gx x
1.19
1
G LTP
L Gx x
1.20
where: μG = absolute viscosity of the gas-phase, kg/m·s
μL = absolute viscosity of the liquid-phase, kg/m·s
The mass flux of the mixture is given by Eq. (1.21).
G LTP TP
V Vm
A
1.21
where: TPm = mass flux, kg/m
2·s
A = cross sectional area of the flow, m2
From the given relations above the pressure drop can be predicted from Eq.
(1.22). Eq. (1.22) neglects the accelerational pressure gradient.
2
2sin
TP TP
TP
TP
f m Lp g L
D
1.22
where: fTP = friction factor, dimensionless
D = diameter of the pipe, m
L = length of the pipe, m
α = angle of the pipe with respect to the horizontal, deg.
17
The friction factor is determined from
16
ReTP
TP
f
1.23
for laminar flow (Re < 2,000), or
1/40.079ReTP TPf
1.23
for turbulent flow (Re > 2,000).
18
CHAPTER 2
REVIEW OF RELATED LITERATURE
2.1. Pyrolysis of Biomass
Pyrolysis is the thermal degradation of organic waste in the absence of oxygen to
produce a carbonaceous char, oil and combustible gases.[25]
Pyrolysis may also be
described as follows. When the drying of a small fuel particle or a zone within a large
particle is completed, the temperature rises and the solid fuel begins to decompose,
releasing volatiles. Since the volatiles flow out of the solid through the pores, external
oxygen cannot penetrate into the particle, and hence the devolatilization is referred to as
the pyrolysis stage.[6]
Unlike combustion in an excess of air, which is highly exothermic
and produces primarily heat and carbon dioxide, pyrolysis of organic material is
analogous to a distillation process and is endothermic.[26]
The high temperatures (900° - 2000°F) and lack of oxygen result in a chemical
breakdown of the waste organic materials into three component streams: (a) a gas
consisting of primarily hydrogen, methane, carbon monoxide, and carbon dioxide, (b) a
“tar” and “oil” that is liquid at room temperature and includes organic chemicals such as
acetic acid, acetone, and methanol, and (c) a “char” consisting of almost pure carbon
plus any inerts and mineral salts that enter the process unit. Residence time, temperature
and pressure can be controlled in the pyrolysis reactor to produce various product
combinations. Most complex organic molecules upon pyrolysis will yield a tar, often
referred to as bitumen, and oil and gas will evolve upon further heating. Both tar and oil
are soluble; they are often referred to as the liquid portion. The residue will be char,
which is often referred to as carbon or “coke.” [26]
The amount of each product produced is dependent on the process conditions,
particularly temperature and heating rate. The process conditions are altered to produce
the desired char, gas or oil end product, with the pyrolysis temperature and heating rate
having the most influence on the product distribution. The heat is supplied by indirect
heating, such as the combustion of the gases or oil, or directly by hot gas transfer.
Pyrolysis has the advantage that the gases of oil product derived from the waste can be
used to provide the fuel for the pyrolysis process itself.[25]
19
2.1.1. Bio-oil
Very high heating rates of about 100°C/s to 1000°C/s at temperatures below
650°C and with rapid quenching, lead to the formation of a mainly liquid product, which
is referred to as fast or flash pyrolysis. Liquid yields up to 70% have been reported for
biomass feedstock using flash pyrolysis. In addition, the carbonaceous char and gas
production are minimized. The primary liquid products of pyrolysis are rapidly quenched
and this prevents breakdown of the products to gases in the hot reactor.[25]
Oils derived from biomass have high oxygen content, of the order of 35% by
weight, due to the content of cellulose, hemicelluloses and lignin in the biomass. Biomass
pyrolysis oils derived from flash pyrolysis processes tend to have a lower viscosity and
consist of a single water/oil phase. The oils are therefore high in water, which markedly
reduces their calorific value. Slow pyrolysis produces liquid products with higher
viscosities which tend to have two phases due to the more extensive degree of secondary
reactions which occur. Pyrolysis oils may contain solid char particles due to carry-over
from the pyrolysis reactor.[25]
The crude pyrolysis liquid is usually dark brown and free flowing with a
distinctive smoky smell. Chemically, it approximates to biomass in elemental
composition and is composed of a very complex mixture of oxygenated hydrocarbons
with an appreciable proportion of water from both the original moisture and reaction
product. Solid char may also be present. The elemental composition of bio-oil resembles
that of biomass rather than that of petroleum oils. The single most abundant bio-oil
component is water.[8]
Bio-oil contains substantial amounts of organic acids (acetic acid and formic
acid). It results in a pH of 2 to 3 and an acid number of 50 to 100 mg KOH/g. Bio-oils
can be corrosive to common construction materials, such as carbon, steel, and aluminum,
due to the presence of these acidic components. The complexity and nature of bio-oil
causes some unusual behavior; specifically, properties that change with time are increase
in viscosity, decrease in volatility, phase separation, and the deposition of gums.[21]
20
2.1.2. Liquid Collection
Liquid collection has long been a major difficulty for researchers. The pyrolysis
vapors have similar properties to cigarette smoke and capture by almost all collection
devices is very inefficient. The product vapors are not true vapors but rather a mist or
fume and are typically present in an inert gas at relatively low concentrations which
increases cooling and condensation problems. They can be characterized as a
combination of true vapors, micron sized droplets and polar molecules bonded with water
vapor molecules. This contributes to the collection problem as the aerosols need to be
impinged onto a surface to permit collection, even after cooling to below the dew point
temperature.[7]
Electrostatic precipitators are effective and are now used by many researchers but
can create problems from the polar nature of the product and arcing of the liquids as they
flow, causing the electrostatic precipitator to short out. Larger scale processing usually
employs some type of quenching or contact with cooled liquid product which is effective.
Careful design is needed to avoid blockage from differential condensation of heavy ends.
The rate of cooling appears to be important. Slow cooling leads to preferential collection
of the lignin derived components which is a viscous liquid which can lead to blockage of
heat exchange equipment and liquid fractionation. Very rapid cooling of the product has
been suggested to be effective as occurs typically in a direct contact quench.[7]
2.1.3. Pyrolysis Gas
The gases produced from biomass waste pyrolysis are mainly carbon dioxide,
carbon monoxide, hydrogen, methane and lower concentrations of other hydrocarbon
gases.[26][25]
The high concentration of carbon dioxide and carbon monoxide is derived
from the oxygenated structures in the original material, such as cellulose, hemicellulose
and lignin. In addition, the gas contains a significant proportion of uncondensed pyrolysis
oils.[25]
21
2.2. Double-Pipe versus Other Types of Condensers
Other condenser/heat exchanger geometry and flow arrangements were reviewed;
and their advantages and disadvantages were compared to the double-pipe condenser.
Reasons for selecting the double-pipe in the initial design are discussed here and also
mentioned through this text. The shell-and-tube and gasketed plate condensers were
considered for future condenser designs as discussed in Chapter 5.
2.2.1. Shell-and-Tube
The shell-and-tube type, shown in Figure 2.1, consists of a large cylindrical shell
inside which there is a bundle of tubes. One fluid stream flows inside the tubes, the other
on the outside of the shell side. Condensation may occur outside or inside the tubes,
depending on the circumstances.[14]
They are perhaps the most common type of heat
exchanger in industrial applications.
Figure 2.1: Shell-and-Tube Condenser[10]
The tubular exchangers are widely used in industry for the following reasons.
They are custom designed for virtually any capacity and operating conditions, such as
from high vacuums to ultra-high pressures (over 100 MPa or 15,000 psig), from
cryogenics to high temperatures (about ll00°C, 2000°F), and any temperature and
pressure differences between the fluids, limited only by the materials of construction.
They can be designed for special operating conditions: vibration, heavy fouling, highly
viscous fluids, erosion, corrosion, toxicity, radioactivity, multicomponent mixtures, and
so on. They are the most versatile exchangers made from a variety of metal and nonmetal
22
materials (graphite, glass, and Teflon) and in sizes from small (0.1 m 2, 1 ft 2) to super-
giant (over 100,000 m 2, 10 6 ft2).[22]
Shell-and-tube heat exchangers require a considerable amount of space, support
structure, capital and installation costs.[22]
For smaller surface area requirements, the
double-pipe is more economical and easier to construct.
2.2.2. Spiral-Tube
Spiral-tube heat exchangers consist of spirally wound coils placed in a shell or
designed as co-axial condensers and con-axial evaporators that are used in refrigeration
systems. The heat transfer coefficient is higher in a spiral tube than in a straight tube.
Spiral-tube heat exchangers are suitable for thermal expansion and clean fluids, since
cleaning is almost impossible.[17]
Bio-oil on the other hand cannot be considered as a clean fluid. As discussed
above, solid char may also be present in the oil[8]
and also lignin derived components
which are viscous liquids which could cause blockage of the condenser.[7]
Compared to
spiral-tube, a double-pipe condenser is easily cleanable because of its simple geometric
construction.
2.2.3. Plate-Fin
Figure 2.2 shows the general form of a plate-fin or simply plate condenser.
Figure 2.2: Plate-Fin Condenser[4]
23
The fluids are separated by flat plates, sometimes there are corrugated fins sandwiched
between the plates. They are often used for low temperature (cryogenics) plants and
where the temperature differences between the streams are small (1 - 5°C). The flow
channels in plate-fin condensers are small and often contain many interruptions to flow.
This can make the channels prone to fouling, which, combined with the fact that they
cannot be mechanically cleaned, means that plate-fin condensers are restricted to clean
fluids.[14]
The same restrictions to the use of spiral-tube are present in plate-fin: they are
both restricted to clean fluid only because they cannot be mechanically cleaned.
2.2.4. Gasketed Plate
Gasketed plate or plate and frame heat exchangers, shown in Figure 2.3, have
several advantages. They are relatively inexpensive and they are easy to dismantle and
clean. The surface area enhancement due to the many corrugations means that a great
deal of surface can be packed into a rather small volume. Moreover, plate and frame heat
exchangers can accommodate a wide range of fluids.[4]
Figure 2.3: Gasketed Plate Heat Exchanger[10]
The design of plate heat exchangers is highly specialized in nature considering the
variety of designs available for the plates and arrangements that possibly suits various
24
duties. Unlike tubular heat exchangers for which design data and methods are easily
available, plate exchanger design continues to be proprietary in nature.[4]
Because of the gasket, they are vulnerable to leakage and hence must be used at
low pressures. The rather small equivalent diameter of the passages makes the pressure
loss relatively high, and the plate and frame heat exchanger may require a substantial
investment in the pumping system, which may make the exchanger cost wise
noncompetitive. Since the flow passages are quite small, strong eddying gives high heat
transfer coefficients, and high local shear which minimizes fouling.[4]
In spite of its many advantages, the gasketed plate was not selected for the initial
design of this study because of the unknown operating pressure of the pyrolysis reactor.
The pressure might be too high for the gasket to withstand or too low to provide the
driving force needed by the volatiles to traverse the condenser. With the double-pipe,
there are no obstructions to constrict the flow of volatiles. Another disadvantage of
gasketed plate is that they are less suitable for condensing duties.[4]
2.2.5. Spiral Plate
With spiral plate heat exchangers the ideal flow conditions and smallest possible
heating surfaces are obtained. The two spiral paths introduce a secondary flow, increasing
the heat transfer and reducing fouling deposits. They are particularly effective in handling
sludge, viscous liquids, and liquids with solids in suspension including slurries. Theses
heat exchangers are quite compact but relatively expensive due to their specialized
fabrication.[17]
Fabrication constraints are the advantages of the double-pipe to a spiral plate
condenser.
2.2.6. Direct Contact
Direct contact condensers are inexpensive and simple to design but have limited
application because the condensate and coolant are mixed. The main advantage of these
condensers, besides their low cost, is that they cannot be fouled and they have very high
heat transfer rates per unit volume.[17]
25
Direct contact condensers were not considered because the volume of the bio-oil
needed to be measured, and employing a direct contact condenser will give difficulty in
measuring the volume since the condensate and cooling medium are mixed.[14]
2.3. Condensers Used in Pyrolysis
2.3.1. Unapumnuk (1999)[24]
The gaseous products of pyrolysis were passed directly through the oil
condensation system. The oil condenser unit was filled with dry ice during the process to
maintain a temperature below 0°C. The oil condenser unit was connected to a glass fiber
filter (Whatman type GF/F filter, diameter 47 mm.). The Millipore filter holder was
controlled at a temperature of 100°C during the experiment.
2.3.2. Mudulodu (2002)[20]
The hot volatiles come out from the reactor top and enter at the top of the
condenser unit. The volatiles passing downward are cooled and subsequently passed
through a liquid collector and a tar filter. The experimental setup is shown in Figure 2.4.
Figure 2.4: Experimental Set-up of Mudulodu (2002)[20]
26
2.3.3. Jih (1982)[16]
The purpose of the condenser (C2) was to condense and collect the tar generated
during the pyrolysis. It was constructed in the same way as the cooler, but its length was
11 inches (279 mm). Prior to each run the inside condenser wall was covered with
aluminum foil and the interior space was loosely filled with steel wool. The amount of tar
collected was determined by subtracting the increasing weight of the foil. In addition to
this, an ice trap (IT) was installed downstream to the condenser (C2) to condense and
collect the remainder of the tar which was not condensed in the condenser. The trap (IT)
was made of Pyrex brand glass (Sargent-Welch Scientific Cat. No. S-82290-A). The trap
was placed in a container which was filled with ice water. Schematic of the experiment
set-up is shown in Figure 2.5.
Figure 2.5: Experimental Set-up of Jih (1982)[16]
2.3.4. Añora (2010)[1]
The condenser was used to separate the condensable gas (bio-oil) and non-
condensable gas (pyrolysis gas) from gas (volatile matter) released during the pyrolysis
reaction. In the condenser, the cooling water absorbed the heat from the gas coming from
27
the distilling flask and resulted to the condensation of bio-oil. Figure 2.6 shows the
experiment set-up used in the study.
Figure 2.6: Experimental Set-up of Añora (2010)
[1]
2.4. Condensation of Mixtures
Heat transfer prediction during condensation of mixtures is more difficult than
during pure vapor condensation for a variety of reasons. For example, with mixtures,
complete or partial condensation can occur depending on whether the coolant
temperature is less than the saturation temperature of the more volatile components.
Along the condenser, as the less volatile components condense out, the concentration of
the more volatile components will increase, and this process creates a vapor temperature
decrease that reduces the driving force for condensation through the condenser. Also, the
presence of different vapor/gas components introduces mass transfer effects that create an
additional thermal resistance that is nonexistent with pure vapors. As a consequence,
condensing heat transfer coefficients of mixtures are less than those of single-component
pure vapors.[22]
Experimental studies show that the presence of noncondensable gases in the vapor
has a detrimental effect on condensation heat transfer. Even small amounts of a
noncondensable gas in the vapor cause significant drops in heat transfer coefficient
during condensation. For example, the presence of less than 1 percent (by mass) of air in
steam can reduce the condensation heat transfer coefficient by more than half.[10]
28
The drastic reduction in the condensation heat transfer coefficient in the presence
of a noncondensable gas can be explained as follows: When the vapor mixed with a
noncondensable gas condenses, only the noncondensable gas remains in the vicinity of
the surface. This gas layer acts as a barrier between the vapor and the surface, and makes
it difficult for the vapor to reach the surface. The vapor now must diffuse through the
noncondensable gas first before reaching the surface, and this reduces the effectiveness of
the condensation process.[10]
2.5. Research Gap
The pyrolysis of marine florae is a relatively new research and is still in its
infancy stage here in the Philippines. There are many existing studies about pyrolysis and
the equipment design and configuration. Examples of these researches are given in
references [7], [8], [16], [20], and [24]. However, the equipment used in a pyrolysis
system is unique to the feedstock and type of pyrolysis process. Thus, the condensing and
liquid collection system described in the studies reviewed above may not be applicable to
the pyrolysis system involving marine florae as feedstock. Hence, the present study
would like to provide preliminary knowledge on the thermal design of the condenser
specific for the developed marine florae reactor developed by Esgana (2011).[13]
29
CHAPTER 3
METHODOLOGY
3.1. Introduction
The condenser development is an iterative process converging towards the
optimum design. The initial design has to be tested under actual operating conditions and
must be evaluated for improvements if necessary. The improved design is again tested
and the process is repeated over again. The present study, however, is limited only to the
first step of the development process which is illustrated in Figure 3.1.
Figure 3.1: Study Flow
30
This study started by calculating the size of the condenser to be able to perform its
specified thermal duties, as discussed later. Then, other parts, pipe fittings, and necessary
equipment that supplement the entirety of the condenser and its operation were selected
mostly by economical basis. The condenser was fabricated when the entire design was
completed. While still on design and fabrication stage of the study, a parallel study
prepared the marine florae feedstock that was used in the pyrolysis experiment. After the
fabrication of the condenser and preparation of the feedstock had been completed, the
study proceeded to the pyrolysis experiment. Here the condenser was tested and
evaluated on its performance, as discussed later. Data regarding the bio-oil and pyrolysis
gas were also collected, since they constitute the fluid that the condenser design was
based on. After all the necessary data had been collected, the initial condenser design was
revised based on the collected data and the size of the condenser was recalculated.
The detailed procedure for the design of the condenser is discussed in Section 3.2.
The feedstock preparation is discussed briefly in Section 3.3. The experiment and data
collection procedures of the present study are discussed in Section 3.5. The reader is
referred to “Design, Fabrication and Test of a Pyrolysis Reactor for Marine Florae” by
Esgana[13]
for the other details of the pyrolysis experiment, i.e. control of pyrolysis
reactor.
3.2. Condenser Design Process
The condenser design was mainly a sizing problem, wherein the heat transfer area
was determined. The tube diameters were selected from standard size pipes, as discussed
in Section 3.2.4. When the suitable tube diameters were selected, the length of the
condenser was solved as discussed below. The following equations used in the design of
the condenser are referred from the equations discussed in Section 1.6. The other
parameters necessary for calculating the heat transfer area are also discussed below.
These parameters are: 1) required heat transfer, 2) convection heat transfer coefficient of
both the cooling water and volatiles, and 3) log mean temperature difference or LMTD.
31
The design flow chart in Figure 3.2[18]
was used as a guide in designing the
condenser but was not strictly followed. The present section is only an overview of the
design process. The complete step-by-step solution is discussed in Appendix A.
Figure 3.2: Condenser Design Flow Chart[18]
3.2.1. Required Heat Transfer
In order to condense the bio-oil, a certain amount of heat must be removed from
the volatiles. To calculate the required heat rejection, the volatiles‟ composition and
properties must be known. However, these data were not available during the initial
design of the condenser. Assumptions were made regarding the composition and
properties of the volatiles to be able to estimate the amount of heat rejection. The
32
condensable component was assumed to be water since literatures support that the most
abundant bio-oil component is water.[8]
The gases produced from biomass waste
pyrolysis are mainly carbon dioxide, carbon monoxide, hydrogen, methane and lower
concentrations of other hydrocarbon gases.[25][26]
The pyrolysis gas was assumed to be
proportions of carbon dioxide, carbon monoxide, hydrogen, and methane. The assumed
mass percentage of each gas component was manipulated to yield the maximum possible
heat transfer requirement.
The mass flow rates of the bio-oil and pyrolysis gas were estimated based on the
results of Añora (2010). Eq. (3.1) and (3.2) were used to solve to mass flow rates, where
ṁbo and ṁpg are the estimated bio-oil and pyrolysis gas mass flow rates, respectively; m
was the design mass capacity of the reactor, which was 5 kg; %bo is the product
percentage of bio-oil, %pg is the product percentage of pyrolysis gas, and t is the
residence time of the pyrolysis experiment by Añora (2010). This procedure assumes that
the average percent composition of bio-oil and pyrolysis gas flow for the entire
experiment remains constant. Details of the calculation are shown in Appendix A. The
mass flow rates computed in Eq. (3.1) and (3.2) were used to estimate the required
amount of heat that must be rejected by the volatiles.
%bo
bo mm
t 3.1
%pg
pg mm
t
3.2
The heat rejections for the bio-oil and pyrolysis gas were estimated as discussed
below. Since the bio-oil was assumed to be water, the heat rejection process would follow
Eq. (3.3).
bo bo i sat fg w sat v,exQ m h h h c T T
3.3
where the enthalpies hi and hsat were obtained from steam tables; hfg and cw are the latent
heat of vaporization and specific heat of water; Tsat is the saturation temperature of water
33
at 1 atm; Tv,ex was the presumed exit temperature of the volatiles. Since the designed
volatile exit temperature from the reactor was 110°C (Esgana 2010), there is
desuperheating of steam from 110°C to 100°C, then latent heat rejection, and sensible
subcooling of liquid water from 100°C to 31°C. The researcher chose to subcool the
volatiles down to 31°C to overestimate size of the condenser. Also, the bio-oil may
contain more volatile components which condense at lower temperatures than the water
content. However, the condensation of these more volatile components was not included
in the calculation because their existence, and thus their thermophysical properties, was
not certain.
The proper method for calculating the maximum theoretical heat transfer is to
subcool the volatiles to the same temperature as the inlet of the cooling water[10]
, that is,
the volatile exit and cooling water inlet temperatures are equal to 30°C, which is the
assumed cooling water temperature. However, the log mean temperature difference in Eq.
(1.10) is indeterminate if the volatile exit and cooling water inlet temperatures are equal
because ΔT2 in Figure 1.5 is zero. To be able to use Eq. (1.10) the volatiles were assigned
an exit temperature of 31°C
Since the pyrolysis gas does not undergo condensation, the heat released from
110°C to 31°C was calculated using Eq. (3.4).
pg pg p v,in v,exQ ym c T T
3.4
where y was the assumed mass fraction of each gas component; cp is the specific heat of
each gas component; Tv,in was the designed inlet temperature of the volatiles.
The total required heat transfer Q in the condenser is then the sum of the heat
rejected by the bio-oil and pyrolysis gas, shown in Eq. (3.5).
bo pgQ Q Q
3.5
The initial design discussed above did not include the centrifugal blower
discussed in Section 3.4 because the researcher expected the volatiles to flow into and the
condenser by natural draft, the same as in the experiment of Añora (2010), since the
34
reactor was an upflow type. However, during the experiment, this was not the case.
During a test run the researcher observed that very little volatiles went in the condenser.
Most of the volatiles went straight up out the reactor feed port. This is explained in more
detail in Section 3.4. A centrifugal blower was installed in the experiment set-up to force
the volatiles to flow to the condenser. This affected the operation of the condenser
because of the increased mass flow rate induced by the blower. The consequences of
installing the blower are explained in more detail in Chapter 4.
3.2.2. Convection Heat Transfer Coefficient
The convection heat transfer coefficient of the volatiles during condensation can
be solved from Eq. (1.12). The equation was modified accordingly, as shown in Eq.
(3.6).[15]
The vapor Reynolds number was found to be less than 35,000, as discussed in
Appendix A.
1/4
0.680.555
3
w w v fg w sat i w
v
w i sat i
ρ ρ ρ h c T T g sinα kh
μ d T -T
3.6
where ρw is the density of water; ρv is the density of steam at 105°C; μw is absolute
viscosity of water; kw is the thermal conductivity of water; cw is the specific heat of water;
hfg is the latent heat of vaporization of water; Tsat is the saturation temperature of water; Ti
is the inner surface temperature of condenser wall; di is the inner diameter of the inner
tube; α is the tilt angle of the condenser. All the properties of water mentioned above
were evaluated at 100°C and 1 atm. The disadvantage of using Eq. (3.6) is that it was
derived from single component condensation and may overestimate the true convection
coefficient of the multicomponent marine florae volatiles. The presence of
noncondensable gases also reduces greatly the convection coefficient.[14]
To compensate
for this shortcoming, the condenser was designed to subcool the volatiles to 31°C, as
discussed in Section 3.2.1.
The condenser was tilted at a 20° angle from the horizontal to enhance
condensation. The effect of the tilt is significant only at low vapor velocities and
optimum at 20°.[14]
The solution for the convection coefficient in Eq. (3.6) is for single
35
component condensation only, which was apparently an incorrect method for solving the
convection heat transfer coefficient of the volatiles. However, the determination of the
convection heat transfer coefficient for gas mixture with noncondensable component is a
very complicated procedure[14]
, and was not possible because of the lack of data on the
properties of the volatiles. To be able to obtain a value for the convection coefficient the
volatiles was treated as a single component water vapor instead of a mixture.
The vapor Reynolds number of the volatiles was determined from Eq. (3.7).
Re
bo pgi vv
v i v
m md m
Aμ πd μ
3.7
where A is the cross sectional area of the inner tube; μv is the absolute viscosity of
volatiles which was assumed to be steam.
The exit velocity of the cooling water was computed using the modified Bernoulli
equation Eq. (3.8) where z1 was the estimated height of the upper reservoir.
2 12v gz
3.8
Since the inner tube diameter was selected to be 1 in., as discussed in Section
3.2.4, the outer tube was chosen based on the Reynolds number. A high Reynolds number
was desirable to attain a high convection heat transfer coefficient, also discussed in
Section 3.2.4. After selecting the size of the outer tube, the Reynolds number of the
cooling water was determined from Eq. (3.9) and was found to be turbulent.
Re w w H
w
w
ρ v D
μ
3.9
where vw was computed from v2 and is discussed in Appendix A; DH is the hydraulic
diameter of the annulus; the cooling water properties, i.e. ρw and μw, were evaluated at
36
30°C. For turbulent flow, the Nusselt number and convection heat transfer coefficient
were computed from Eq. (3.10) and (3.11), respectively.[15]
0.8 0.4Nu 0.023 Re Prw w w
3.10
Nuw ww
H
kh
D
3.11
where Prw and kw are the Prandtl number and thermal conductivity of water at 30°C,
respectively.
3.2.3. Logarithmic Mean Temperature Difference
Based on the total heat rejection computed from Eq. (3.5), the exit temperature of
the cooling water was determined using Eq. (3.12). The inlet temperature of the cooling
water was approximated as 30°C.
w,ex w,in
w w
QT T
ρ Vc
3.12
where Tw,in was the assumed inlet temperature of the cooling water which was 30°C; V is
the volume flow rate of cooling water.
The log mean temperature difference was calculated from Eq. (1.10), repeated in
Eq. (3.13). The temperature profile is similar to Figure 1.5.
1 2
1 2lnln
v,in w,ex v,ex w,in
lm
v,in w,ex
v,ex w,in
T T T TT TT
T T T T
T T
3.13
3.2.4. Heat Transfer Area
A 1-in. diameter pipe was selected as the inner tube because the diameter of the
reactor gas-exit-pipe was also a 1 in. diameter pipe. The diameter of the outer tube was
determined based on the Reynolds number of the cooling water, as previously discussed.
37
If the diameter of the outer tube is small the Reynolds number is high, and opposite for
large diameter tubes. A high Reynolds number was desirable to attain a high convection
heat transfer coefficient, as discussed in the Section 3.2.2. The diameter of the outer tube
was restricted to standard size GI pipes only. GI pipe was chosen as outer tube because
they are cheaper than stainless steel pipes but more rigid that aluminum pipes; the
research was not concerned with the heat transfer and corrosion in the outer tube.
After selecting the inner and outer tube diameters, the length of the condenser was
calculated using Eq. (3.14). The aluminum condenser length was 78.1 cm, and the
stainless condenser was 99.9 cm. The details of the calculations are explained in
Appendix A.
ln1 1
Δ 2
o i
lm i v t o w
d dQL
π T d h k d h
3.14
3.3. Marine Florae Collection and Preparation
The collected marine florae were segregated according to the type and then sun-
dried. Once dried, the different types of marine florae were then pulverized to prepare for
pelletizing. Some of the pulverized marine florae were mixed with a water-cornstarch
binder. The proportion was 80% marine florae and 20% binder by weight. The binder
was 40% water and 60% cornstarch by weight. Pure pulverized marine florae, i.e. without
binder, were also pelletized. Raw pulverized (not pelletized) marine florae were also used
as feedstock for the pyrolysis process. There were eight total different types of marine
florae feedstock which are listed in the Table 1.1 in Section 1.5.2.
3.4. Installation of Centrifugal Blower
During a test run the researcher observed that very little volatiles went out the
gas-exit-valve of the condenser. Because of this, the test run was ended early. When the
reactor feed port was opened it was seen that there were plenty of gases trapped inside the
reactor. These gases escaped out to the atmosphere through the feed port.
38
Figure 3.3 shows the gases escaping out the reactor. The pressure inside the reactor might
not have been enough to provide the draft for the volatiles to flow through the designed
gas-exit-pipe and to the condenser. Because of the orientation and relatively small
opening of the gas-exit-pipe the volatiles needed an external force to direct their flow.
Figure 3.3: Gas Escaping through the Feed Port of the Reactor
A centrifugal blower was installed to direct the flow of the volatiles to the condenser. The
said blower was chosen because it was easily retrofitted to match the dimensions of the
reactor gas-exit-pipe and condenser inlet. The blower has an indicated speed of 3,000-
3,600 rpm at 50-60 Hz. The suction and discharge diameters were 4 in. and 2 in.,
respectively. The actual rpm and air velocity before retrofitting were measured using a
digital tachometer and analog velometer, respectively, as illustrated in Figure 3.4. The
rpm was 3,305 and the air velocity was approximately 1,450 fpm or 7.368 m/s.
Figure 3.4: RPM and Air Velocity Measurement
39
The suction and discharge ports were both retrofitted to fit the 1 in. diameter of the
reactor gas-exit pipe and the condenser inner tube. Due to the installation of the blower
the condenser‟s designed tilt angle of 20° was not realized. The actual tilt angle is
discussed in Section 4.1. The retrofitted centrifugal blower is shown in Figure 3.5.
Figure 3.5: Retrofitted Centrifugal Blower
3.5. Experiment Set-up and Procedure
The present study was done simultaneously with Esgana‟s research[13]
, that was
about reactor design and performance evaluation. The reader is referred to “Design,
Fabrication and Test of a Pyrolysis Reactor for Marine Florae” by Esgana (2011) for the
procedures of when to load and unload the feedstock, and control of temperatures inside
the reactor. Figure 3.6 illustrates the schematic diagram of the experiment set-up.
Figure 3.6: Schematic of Experiment Set-up
40
The volatiles were sucked out of the reactor by means of the centrifugal blower. Before
the volatiles enter the condenser its static pressure and temperature were measured. Upon
leaving the condenser its temperature was again measured. The adapter allowed
separation of flow of the bio-oil and the pyrolysis gas. The bio-oil was collected in a
beaker and the exit velocity of the pyrolysis gas was measured. Figure 3.7 shows the
actual experiment set-up without the pressure and velocity measuring equipments.
Figure 3.7: Actual Experiment Set-up without Manometer
The experiments were conducted at the University of San Carlos Mechanical
Engineering Laboratory. In the present study an experiment involving one type of
feedstock and either of the two condensers is called a „run‟. The number of runs that were
conducted in a single day was limited to the length of the duration of one run which was
4 to 6 hours, depending on the type of feedstock. A maximum of two runs were
conducted in one day because the experiments were conducted during school hours only.
41
In some days only one run was conducted. The runs of the experiment are listed in Table
3.1.
Table 3.1: Experiment Runs
Run No. Condenser Date Feedstock
A1 Aluminum 2/24/11 Pure Green Pellets
A2 Aluminum 2/25/11 Pure Red Pellets
A3 Aluminum 2/25/11 Green Raw
A4 Aluminum 3/3/11 Pure Green Pellets (2nd
)
A5 Aluminum 3/3/11 Red Raw
A6 Aluminum 3/5/11 Pure Seagrass Pellets
A7 Aluminum 3/6/11 Pure Brown Pellets
A8 Aluminum 3/7/11 Seagrass with Binder
S1 Stainless 2/19/11 Brown with Binder
S2 Stainless 2/21/11 Seagrass with Binder
S3 Stainless 2/22/11 Red Raw
S4 Stainless 2/22/11 Pure Red Pellets
S5 Stainless 2/23/11 Pure Green Pellets
S6 Stainless 2/23/11 Green Raw
S7 Stainless 3/2/11 Pure Brown Pellets
S8 Stainless 3/2/11 Pure Seagrass Pellets
3.5.1. Equipment Preparation
The condenser and blower were cleaned after a single day of experimentation to
prepare for the next experiment. A single day had either one or two runs. When two runs
were done the condenser and blower were cleaned only after the last run. The blower and
condenser were dismantled from the pyrolysis set-up and rinsed with tap water, then
wiped dry with cloth. After drying, the blower and condenser was reattached to the
pyrolysis reactor.
As designed, the condenser was installed at an angle with the horizontal. The
designed tilt angle of 20° was not realized in the actual experiment set-up because of the
installed centrifugal blower. The actual tilt angle is discussed in Section 4.1. The
42
condenser was also insulated with rock wool to minimize heat exchange with the
environment. Figure 3.8 shows the insulated condenser.
Figure 3.8: Insulated Condenser
The upper reservoir was filled with tap water and the water was allowed to flow
through the water side of the condenser and to the lower reservoir. When the lower
reservoir was filled to a sufficient level the water pump was turned on to recirculate the
water back to the upper reservoir.
3.5.2. Cooling Water Flow Calibration
The flow of the cooling water was controlled with the valve shown in Figure 3.9.
Figure 3.9: Installed Condenser
For a certain valve opening, a steel can placed about the same height as the lower
reservoir, was filled with the water leaving the exit hose. The time at which it took to fill
the steel can was recorded and the amount of water in the steel can was weighed. The
43
weight of the water divided by the time it took to fill the steel can was the mass flow rate
of the cooling water for a certain valve opening.
The valve openings that were measured were „fully open‟, „one valve turn‟, „two
valve turns‟, „three valve turns‟, and „four valve turns‟. Five valve turns was not included
in the calibration because the valve was nearly fully closed. Five trials were done for each
valve openings mentioned earlier. Results of the calibration are tabulated in Appendix C.
The flow rate of the cooling water was varied throughout the experiment to determine if
it significantly affected the heat transfer. The result of this trial is discussed in Section
4.2.2.
3.5.3. Fluid Temperature Measurement
The inlet and exit temperatures of the volatiles and cooling water were measured
with the digital thermocouple datalogger, shown in Figure 3.10.
Figure 3.10: Thermocouple Datalogger
Even though the temperature rise of the cooling water was predicted to be very small,
approximately 0.546°C (refer to Appendix A), the researcher still decided to measure
both inlet and exit temperatures for the fact the actual experiment conditions could vary
from the calculations because of the numerous assumptions made.
44
Figure 3.11 shows the placement of the thermocouple probes in the condenser. Appendix
B shows the exact positioning of the thermocouple probes in the two condensers.
Figure 3.11: Condenser with Thermocouple Probes
3.5.4. Periodic Oil Collection and Measurement
In the original methodology only the total collected volume of the bio-oil was
supposed to be recorded. This was done for the first four runs of the stainless steel
condenser which were runs S1, S2, S3, and S4. However, the researcher observed that the
rate of bio-oil yield was not constant. The bio-oil yield increased whenever the blower
was turned on, as discussed in Section 4.4.1. Because of this observation the bio-oil yield
was measured periodically. Figure 3.12 shows a simple illustration of how the bio-oil was
collected and measured.
Figure 3.12: Bio-oil Collection and Storage
The volume of the bio-oil collected in the beaker was recorded using a graduated cylinder
for better accuracy. This was done every 15 minutes starting from the time the first drops
of bio-oil were observed. After recording the volume of oil in each 15-minute interval it
45
was transferred to a glass bottle for storage. Once transferred, the weight of the collected
oil was determined using a digital weighing scale.
3.5.5. Static Pressure Measurement
The static pressure of at the inlet of the condenser was measured using an inclined
manometer as shown in the Figure 3.13. The manometer was inclined at 30° with respect
to the horizontal. The inclination angle was positioned by using a 30° x 60° triangle and a
hose filled with water used as a level gage shown in Figure 3.14.
Figure 3.13: Static Pressure Measurement Set-up
Figure 3.14: Inclination Positioning Instruments
The base of the triangle was in lined with the horizontal by means of the water-hose level
gage. The side of the manometer was then inclined until it was parallel with the
hypotenuse of the triangle.
a) 30° x 60° Triangle b) Water-Hose Level Gage
46
The static pressure was measured for the Aluminum condenser only since it has
more potential of being used as condenser material because it was easier to clean than the
stainless condenser. Also, more bio-oil can be collected since less stick to the walls of the
aluminum condenser than in the stainless condenser. Furthermore, the material roughness
was not considered in the calculation of the pressure drop. With the static pressure known
the gas density could be solved from the ideal gas law as shown in Eq. (3.15). The gas
density was necessary in calculating the heat transfer, as explained in Section 3.6.2.
p
RT
3.15
Measurement of the static pressure required attaching additional pipe fittings to
the condenser. These pipe fittings were made of galvanized iron, and the bio-oil sticks to
them which could decrease the amount of bio-oil collected in the beaker, thus, only runs
A4 and A5 were subjected to static pressure measurements to minimize bio-oil loss. Only
the inlet static pressure was measured because only one manometer was available.
Appendix B shows the actual positions of the pressure taps with dimensions.
Measurement of inlet and exit static pressure must be made simultaneously because the
volatile flow rate was not steady. The unsteady flow was indicated by the fluctuating
temperature; the reactor temperature profile also varied with time which means the
devolatilization was not constant. The researcher also attempted to measure the
differential pressure; however, there was very little change in water column height, which
was unreadable. The calculations for the exit static pressure are discussed in Section
3.6.2. The static pressure was measured only when the blower was turned on since there
was no observed change in water column height when the blower was turned off.
3.5.6. Gas Velocity Measurement
The pyrolysis gas velocity at the gas-exit-valve of the adapter was measured when
the blower was turned on using an analog velometer. It was observed that when the gas-
exit-valve was fully opened, the exit temperature of the pyrolysis gas was relatively high,
as high as 60°C. When the gas valve was slightly closed, about 45° angle of the lever; the
47
gas exit temperature was low, sometimes as low as the cooling water inlet temperature.
This meant that when the blower was turned on, the flow of the volatiles could be varied
by varying the opening of the gas-exit-valve. The purpose of varying the flow was to try
different flow velocities and analyze the optimum velocity of the flow. The analysis is
discussed in Section 4.3 and 4.7. Figure 3.15 shows the gas-exit-valve positions.
Figure 3.15: Gas-Exit-Valve Positions
With the exit velocity of the gas measured, the gas velocity inside the inner tube
and mass flow rate could be calculated from Eq. (3.16) and (3.17), respectively.
ex exA vv
A
3.16
m Av
3.17
where v is the gas velocity inside the inner tube; ṁ is the mass flow rate; Aex is the cross
sectional area of the gas-exit-valve of the condenser; vex is the velocity measured by the
velometer; ρ is the gas density; A is the cross sectional area of the inner tube. The mass
flow rate was used in evaluating the performance of the condenser which is discussed in
Section 3.6. The gas velocities for runs A4 and A5 only, the same runs for the static
pressure measurement, were measured. The gas velocity when the blower was off was
not measured because the velocity was too low for the velometer to read.
c) Closed b) Slightly closed a) Full open
48
3.5.7. Gas Collection for Gas Chromatography
Pyrolysis gas samples were collected using uro-bags and were sent to the
University of San Carlos Chemical Engineering Laboratory for chromatography. The
Shimadzu GC8A gas chromatograph apparatus was used in the analysis of the gas
composition. The gases that it is able to analyze, however, were limited to carbon dioxide
and methane only. The carbon dioxide and methane content of the pyrolysis gas are
tabulated in Section 4.5. A picture of an uro-bag filled with pyrolysis gas is shown in
Figure 3.16.
Figure 3.16: Uro-bag filled with Pyrolysis Gas
3.6. Condenser Evaluation
The effectiveness of the condenser was not calculated, as originally planned,
because of the inconsistencies in the cooling water temperature readings, discussed in
Section 4.2.2. A homogeneous two-phase model was used to estimate the pressure drop
and the actual heat transferred.
3.6.1. Cleanability
It was observed by Añora (2010) in his experiment that the early condensation of
bio-oil while still inside the reactor resulted in deposits of bio-oil in the reactor walls. In
the present study, both the aluminum and stainless condensers were visually inspected for
bio-oil deposits in its walls. The condenser material with fewer deposits and easier to
clean was identified.
49
3.6.2. Pressure Drop
A homogeneous two-phase flow model proposed by reference [14] was used to
calculate the pressure drop in the inner tube. Since only runs A4 and A5 had data on the
static pressure and gas velocity, the pressure drop was solved for runs A4 and A5 only.
The static pressure and gas velocity measurements were taken only once for each gas-
exit-valve opening (full open and slightly closed) in each run (A4 and A5); the
measurements were not taken for the entire run. Hence, the calculated pressure drop is
valid only for the short time-duration that the static pressure and gas velocity
measurements were taken. The measurements for runs A4 and A5 were taken at 1:43:00
to 1:46:50 and 0:15:30 to 0:19:30, respectively. The actual heat transferred was also
estimated based on the same instance when the static pressure and gas velocity were
measured, as discussed in Section 3.6.3.
The homogeneous two-phase flow model assumes that the flow velocity of the
pyrolysis gas is the same as the bio-oil. This assumption was made because the researcher
had no way of knowing the actual velocity of the bio-oil, only its volume flow rate; only
the velocity of the pyrolysis gas was measured in the experiment, refer to section 3.5.6.
Thus, the velocity of the two-phase flow is
TP G Lv v v
3.18
where vTP is the two-phase flow velocity that was used in the calculation of the pressure
drop; vG was the estimated pyrolysis gas velocity; vL is the velocity of the liquid bio-oil,
which in this case is assumed to be equal to the gas velocity.
The properties of the bio-oil were again assumed to be the same as water because
some of the properties, which are specific heat and thermal conductivity, were not
determined. For consistency, the properties of water were used throughout the
calculation. The properties of water were evaluated at the average temperature of the
flow. It was also discovered in the calculation that using the actual density of the bio-oil
did not have a significant effect on the numerical value of the pressure drop. The
properties of the pyrolysis gas were determined based on its composition tabulated in
Table 4.6 in Section 4.5.1. Run A4 had a Pure Green Pellet feedstock that has a
50
composition of 92.08% CO2 and 7.92% CH4; run A5 used a Red Raw feedstock that has a
composition of 75% CO2 and 23% CH4. All gas properties used in the calculation were
evaluated at the average temperature of the flow. The complete solution of the
determination of the pressure drop is shown in Appendix G.
The densities of each gas component, which are CO2 and CH4, were calculated
from the ideal gas equation shown in Eq. (3.19).
p
RT
3.19
where ρ is the gas density, p is the static pressure of the flow, R is the gas constant, and T
is the absolute temperature of the gas. The symbols ρ1 and ρ2 represent the gas densities
at the inlet and exit of the condenser, respectively. For ρ1, p1 is the inlet static pressure
that was measured in the experiment. Since there was no data on the exit static pressure
p2, p1 was used to solve ρ2. After the pressure drop was solved, the exit static pressure p2
was determined and inserted back to the original solution of ρ2. The new ρ2 was then used
to re-compute the pressure drop in an iterative manner until the solution converges.
Compressibility factors of gas components were not included in Eq. (3.19) because the
compressibility factors for both CO2 and CH4 were very close to unity at operating
conditions of the experiment.[11]
The absolute viscosity of each gas component was determined from gas property
tables. The pyrolysis gas was treated as a homogeneous mixture of CO2 and CH4, thus,
the properties of the gas mixture were calculated from Eq. (3.20) and (3.21).
2 2 4 4G CO CO CH CHy y
3.20
2 2 4 4G CO CO CH CHy y
3.21
where ρG and μG are the density and absolute viscosity of the gas mixture (pyrolysis gas),
respectively; 2COy and
4CHy are the mass fractions of each gas component; 2CO and
4CH are the densities of each gas component computed from Eq. (3.19); 2CO and
4CH
are the absolute viscosities of each gas component.
51
The homogeneous two-phase model requires the values of the volume flow rates
of the bio-oil and pyrolysis gas, which were calculated from Eq. (3.22) and (3.23),
respectively.
15 min
boL
VV
3.22
G TPV Av
3.23
where LV and GV are the volume flow rates of the bio-oil and pyrolysis gas, respectively;
Vbo is the volume of bio-oil collected in the 15-minute period corresponding to the time
when the static pressure and velocity measurements were taken; vTP is the computed
velocity of the pyrolysis gas inside the condenser; A is the flow area of the pyrolysis gas.
Next, the void fraction and the quality of the two-phase flow were determined
from Eq. (3.24) and (3.25), respectively.
GG
L G
V
V V
3.24
1
GG
L
GG G
L
x
3.25
where εG and x are the void fraction and quality, respectively; LV and GV
are the volume
flow rates of the bio-oil and pyrolysis gas, respectively; ρL and ρG are the densities of the
bio-oil and pyrolysis gas, respectively.
52
The quality is then used to calculate the two-phase density and absolute viscosity
as shown in Eq. (3.26) and (3.27), respectively.
1
G LTP
L Gx x
3.26
1
G LTP
L Gx x
3.27
where ρTP and μTP are the two-phase density and absolute viscosity, respectively; x is the
quality; ρL and ρG are the densities of the bio-oil and pyrolysis gas, respectively; μL and
μG are the absolute viscosities of the bio-oil and pyrolysis gas, respectively. The mass
flux TPm was then determined from using Eq. (3.28).
L GTP TP
V Vm
A
3.28
Reynolds number ReTP was computed from Eq. (3.29) to determine if the flow
was laminar or turbulent.
Re TP i
TP
TP
m d
3.29
where TPm is the mass flux; di is the inner diameter of the inner tube; μTP is the two-phase
absolute viscosity. In two-phase flow, Re < 2000 is laminar and Re > 2000 is
turbulent.[14]
Eq. (3.30) and (3.31) were used to solve for the friction factors for laminar
and turbulent flow, respectively.
16
ReTP
TP
f
3.30
1/40.079ReTP TPf
3.31
53
The pressure drop was calculated from Eq. (3.32). Then, the exit pressure was
determined by subtracting the pressure drop from the inlet static pressure measured
during the experiment, as shown in Eq. (3.33).
2
2sin
TP TP
TP
i TP
f m Lp g L
d
3.32
2 1p p p
3.33
where Δp is the pressure drop; fTP is the friction factor based on either laminar or
turbulent flow; L is the distance between the two pressure taps; α is the tilt angle of the
condenser with respect to the horizontal; g is the acceleration due to gravity; ρTP and TPm
are the density and mass flux of the two-phase mixture, respectively; p2 is the exit
pressure; p1 is the inlet pressure.
3.6.3. Actual Heat Transferred
The actual heat transferred in runs A4 and A5 were estimated based on the same
instance when the static pressure and gas velocity were measured, when the blower was
turned on. The reason is that the mass flow rate of the volatiles can be estimated only
when the blower was turned on. Also, the density of the pyrolysis gas was determinable
only when the static pressure of the flow was known. These statements were also
discussed in Section 3.6.2 to solve for the pressure drop. The complete solution of the
determination of the actual heat transferred is shown in Appendix G.
The specific heats at constant pressure of the individual gas components were
determined from gas property tables as a function of temperature only.
[11] Then the
specific heat of the pyrolysis gas was calculated using Eq. (3.34).
2 2 4 4G CO CO CH CHc y c y c
3.34
54
where cG is the specific heat at constant pressure of the pyrolysis gas; 2COc and
4CHc are
the specific heats of each gas component; 2COy and
4CHy are mass fractions of each gas
component. Both the inlet and exit specific heats were calculated from Eq. (3.34), then,
the average specific heat was determined. The specific heat of the bio-oil (assumed as
water) was also evaluated at the average flow temperature. The specific heat of the two-
phase mixture was then calculated using Eq. (3.35).
1
G LTP
L G
c cc
xc x c
3.35
where cTP is the specific heat of the two-phase mixture; cG and cL are the specific heats of
the pyrolysis gas and bio-oil, respectively; x is the quality. The heat transferred is then
calculated using Eq. (3.36).
TP TPQ m Ac T
3.36
where Q is the heat transferred; cTP is the specific heat of the two-phase mixture; ΔT is
the change in temperature of the volatiles that was measured by thermocouples; A is the
cross-sectional area of the flow; TPm is the mass flux.
3.7. Recalculation of the Double-Pipe Condenser Length
Having determined the components of the pyrolysis gas and the amount of bio-oil
extracted, the double-pipe condenser length was recalculated using an improved design
methodology. This design method also uses the homogeneous two-phase flow model that
was used in Section 3.6.2. Data from the experiment, which were not available to the
initial condenser design, were incorporated in the recalculation of the length. These data
are gas velocity, rate of bio-oil yield, pyrolysis gas components, static pressure, and
cooling water flow rate. The condenser length was recalculated while keeping the pipe
diameters constant. The steps of the recalculation are discussed in this section, and the
55
complete solution with numerical values based on the experiment results is presented in
Appendix H.
Unlike the initial condenser design discussed in Section 3.2, the condenser was
divided into three zones: desuperheating zone, condensing zone, and subcooling zone.
Refer to Figure 3.17.
Figure 3.17: Temperature Profile
For the present study, the operating temperature of the desuperheating zone is from
110°C to 100°C since the designed volatile exit temperature from the reactor is 110°C[13]
and water condenses at 100°C. A volatile inlet temperature of 110 °C, however, assumes
that there is no heat loss in the gas-exit-pipe of the reactor and that the temperature of the
volatiles leaving the reactor is equal to the temperature at Layer A of the reactor.[13]
The
condensing zone is at a constant 100°C and the subcooling zone is from 100°C to 31°C.
The researcher chose to subcool the volatiles down to 31°C for the same reasons
discussed in Section 3.2.1. The calculation discussed in the present Section was done for
the six marine florae feedstock whose compositions were analyzed, as shown in Table 4.6
in Section 4.5.
3.7.1. Properties of Bio-oil and Pyrolysis Gas
First, the properties of the bio-oil and pyrolysis gas were determined for the
desired operating temperatures and pressure. Since the thermophysical properties of the
56
bio-oil were still not known, the bio-oil was again assumed as water in the calculations
presented in Appendix H. However, if complete data on the thermophysical properties of
the bio-oil is available they should be used in the calculation in place of the water
properties. The bio-oil properties that were needed in the solution are listed in Table 3.2.
These properties were obtained directly from property tables of water, e.g. steam tables.
Table 3.2: Necessary Bio-oil Properties
Condition Properties
Superheated steam at 110°C specific enthalpy, absolute viscosity,
density, thermal conductivity
Saturated steam at 100°C specific enthalpy, absolute viscosity,
density, thermal conductivity
Saturated water at 100°C specific heat, absolute viscosity, density,
thermal conductivity
Subcooled water at 31°C specific heat, absolute viscosity, density,
thermal conductivity
The properties of the pyrolysis gas were determined from the properties of the
individual gas components. The result of the gas chromatograph showed that these
components are CO2 and CH4. However, other gases might also be present, but was not
detected by the type of gas chromatograph equipment used in this study. The properties
of CO2 and CH4 that were needed in the solution are listed in Table 3.3.
Table 3.3: Necessary Gas Properties
Condition Properties
at 110°C specific heat at constant pressure, absolute
viscosity, density, thermal conductivity
at 100°C specific heat at constant pressure, absolute
viscosity, density, thermal conductivity
at 31°C specific heat at constant pressure, absolute
viscosity, density, thermal conductivity
The specific heat, absolute viscosity, and thermal conductivity were obtained from gas
property tables as a function of temperature only. Their variation with pressure was not
57
considered because of the small pressure gradient of the flow which is discussed later.
The density was calculated using Eq. (3.20) in Section 3.6.2. The pyrolysis gas was
treated as a homogeneous gas mixture and its properties as a whole were calculated using
equations similar to Eq. (3.21) and (3.22) in Section 3.6.2.
3.7.2. Mass Flux
Eq. (3.22) and (3.23) in Section 3.6.2 were used to calculate the volume flow rate
of the bio-oil and pyrolysis gas, respectively; the void fraction, quality, and mass flux
were calculated using Eq. (3.24), (3.25), and (3.28), respectively. The volume flow rate
of the bio-oil was estimated based on the 15-minute sampling rate of the collected bio-oil
volume discussed in Section 3.5.4. In the calculation of the volume flow rate of bio-oil in
Eq. (3.22), the volume of the bio-oil was set to the highest recorded volume of bio-oil in a
15-minute duration which was 80 ml; see Figure F.8 in Appendix F. This is equivalent to
a volume flow rate of 5.33 ml/min. It is evident from the experiment results shown in
Appendix F that the volume of bio-oil collected vary in every 15-minute interval
sampling rate. If the condenser is designed to condense the highest amount of bio-oil then
it is certain to condense the lesser amounts. The value of the velocity that was used to
calculate the volume flow rate of the pyrolysis gas was the actual velocity measured in
the experiment.
The flow parameters discussed above were all calculated based on the actual
experiment condition, wherein the condenser seemed to be just a subcooling heat
exchanger because of undesired condensation of the bio-oil prior to the condenser inlet.
This is discussed further in Section 4.6. The volume flow rates and void fraction may be
different in the desuperheating and condensing zones but the mass flux and quality must
be constant to satisfy the principle of conservation of mass. Disregarding the small
amount of oil that sticks to the condenser walls, the mass leaving the condenser must be
equal to the mass entering the condenser. That is, the mass of the volatiles that leave the
subcooling zone is equal to the mass that enter the subcooling zone, which is the same
mass that leave the condensing zone, and so on.
58
The statement above is illustrated in Figure 3.18 where m represents the mass of the
volatiles.
Figure 3.18: Conservation of Mass in the Condenser
However, in reality the mass entering the condenser is greater than the mass
leaving the condenser because of the small amounts of bio-oil that adhere to the walls of
the condenser and accumulate over time. More about this oil is discussed in Section 4.4.
The consequence of the assumption that the mass flux is constant is that the calculation of
the total required heat transfer will be overestimated. But since the researcher did not
have the means to determine the change in mass flux, the calculations were done with
constant mass flux.
3.7.3. Required Heat Transfer
In the desuperheating zone, the equation used in the calculation of the required
heat transfer was slightly different from that used in Section 3.6.3. In the case of
superheated steam the specific enthalpy must be used instead of the specific heat in the
calculation of heat transfer as shown in Eq. (3.37). The heat released by the pyrolysis gas
was calculated using Eq. (3.38).
1L TPQ m x A h
3.37
G TP GQ m xAc T
3.38
where QL is the amount of heat released by bio-oil (assumed as steam) during
desuperheating from 110°C to 100°C; QG is the heat released by the gas from 110°C to
100°C; TPm and x are the mass flux and quality, respectively, that were determined from
59
Section 3.7.2; A is the cross sectional area of the inner tube; Δh is the change in specific
enthalpy of the bio-oil from 110°C to 100°C; ΔT is simply the difference between 110°C
and 100°C. The total heat released by the volatiles in the desuperheating zone, therefore,
is
des L GQ Q Q
3.39
In the condensing zone, only the heat released by the bio-oil during condensation
was calculated as shown in Eq. (3.40) where hfg is the latent heat of vaporization of water
at 1 atm.
1con TP fgQ m x Ah
3.40
In the subcooling zone, the heat released both by the bio-oil and pyrolysis gas is
calculated using Eq. (3.41).
sub TP TPQ m Ac T
3.41
where cTP is the specific heat of the two-phase mixture calculated from Eq. (3.35) in
Section 3.6.3; ΔT is the temperature difference between 100°C and 31°C.
3.7.4. Logarithmic Mean Temperature Difference
The first step in calculating the LMTD in each zone of the condenser was to
determine the inlet and exit temperatures of the cooling water in each zone. These
temperatures are indicated in Figure 3.17 as Tw1, Tw2, Tw3, and Tw4. Tw1 and Tw2 are the
inlet and exit temperatures in the subcooling zone, respectively; Tw2 is also the inlet
temperature in the condensing zone and Tw3 is the exit temperature; Tw3 is also the inlet
temperature in the desuperheating zone and Tw4 is the exit temperature. In the calculations
Tw1 was set to 30°C. Tw2, Tw3, and Tw4 were calculated using equations similar to Eq.
(3.12) in Section 3.2.3. Afterwards, the LMTD in the desuperheating, condensing, and
subcooling zones were calculated using equations similar to Eq. (3.13). The actual
60
formulas used in the calculations together with the complete solutions are presented in
Appendix H.
3.7.5. Convection Heat Transfer Coefficients
The formulas for calculating convection heat transfer coefficients depend mainly
on the value of the Reynolds number. The Reynolds number of the volatiles in each zone
was determined in order to decide the most applicable equations for solving the
convection heat transfer coefficient of the volatiles in each zone. The convection
coefficient in each zone was calculated separately because of the different heat exchange
duties of each zone. The properties of the bio-oil in each zone differ because it undergoes
phase change and the different phases have different convection coefficient. The
convection heat transfer coefficient of the cooling water was calculated based on the
actual flow rate measured from the experiment. The formulas used in the calculation of
the Reynolds number and convection heat transfer coefficient were already presented in
Sections 1.6 and 3.2. The properties used in the calculation of the convection heat
transfer coefficients of the volatiles, i.e. kTP, μTP, and ρTP, were calculated using equations
similar Eq. (3.26) and (3.27) in Section 3.6.2. The details of the calculation of the
convection heat transfer coefficients are presented in Appendix H.
In the condensing zone, it was assumed in the calculations that the superheated
bio-oil was the only medium, that is, the pyrolysis gas was neglected. This was done to
simplify the calculations. The heat exchange duty of decreasing the pyrolysis gas
temperature below 100°C was assigned to the subcooling zone. This approach was also
done in the calculation of the pressure drop in Section 3.7.7.
3.7.6. Length of the Condenser
In the analysis presented in this section, the condenser is actually divided into
three heat exchangers: 1) desuperheating, 2) condensing, and 3) subcooling heat
exchanger. This concept was explained at the beginning of this section as three condenser
zones. The lengths of each zone were solved individually and then totaled to obtain the
length of the entire condenser. Therefore, there were three equations used to solve the
length of each zone. These equations are similar to Eq. (3.14) in Section 3.2.4. The only
61
variations are the different values of Q, ΔTlm, and hv in each zone. The total length of the
condenser was calculated using Eq. (3.42), where Ldes, Lcon, and Lsub are the lengths of the
desuperheating, condensing, and subcooling zones, respectively.
des con subL L L L
3.42
3.7.7. Pressure Drop
Since it was observed from the test run, as discussed in Section 3.4, that the
blower is essential equipment in the pyrolysis set-up, the pressure drop inside the
condenser must be estimated so that the appropriate size of the blower can be selected.
The same procedure and equations presented in Section 3.6.2 for calculating the pressure
drop were used. Different flow conditions were solved and the results are compared in
Section 4.7. The details of all the calculation presented here in Section 3.7 are presented
in Appendix H.
62
CHAPTER 4
RESULTS AND DISCUSSION
4.1. Designed and Fabricated Double-Pipe Condenser
The designed condenser lengths, indicated by the symbol L in Figure 4.1, were
78.1 cm and 99.9 cm for aluminum and stainless condenser, respectively. However, due
to fabrication errors the actual fabricated lengths were 52 cm and 61 cm for aluminum
and stainless condenser, respectively. The full detail on the designed condenser is shown
in Appendix B. TC1, TC2, TC3, and TC4 in Figure 4.1 indicate the slots where the
thermocouple probes were inserted. TC1 and TC2 measured the inlet and exit
temperatures, respectively, of the volatiles. TC3 and TC4 measured the inlet and exit
temperatures, respectively, of the cooling water. Refer to Appendix B for the exact
positions of the thermocouple probes.
Figure 4.1: Condenser Length
The designed tilt angle of the condenser was also not realized because of the
orientation of the blower with respect to the reactor. The actual tilt angle during the
experiment was 25° with respect to the horizontal as shown in Figure 4.2. The procedure
for measuring the angle is presented in Appendix B.
Figure 4.2: Condenser Tilt Angle
63
4.2. Temperature of Condenser Fluids
4.2.1. Temperature of Volatiles
The inlet temperature of the volatiles in the condenser was much lower than the
temperature recorded in Layer A of the reactor[13]
, shown in Figure 4.3. This meant that
there was heat rejection that caused a temperature drop between the reactor and
condenser. The temperature drop was due to the relatively long distance that the volatiles
had to travel before getting to the condenser inlet. Because the gas-exit-pipe of the
reactor leading to the condenser was initially at a lower temperature than the volatiles in
the reactor, the volatiles gave off heat as they pass through the pipe. The gas-exit pipe
was insulated, so, theoretically, there had to be a certain time when the pipe temperature
will attain thermal equilibrium with the volatiles. However, still before the condenser
inlet was the blower which was not insulated. Continuous heat rejection to the
environment occurred in the blower which led to the large temperature drop and
condensation of bio-oil in the blower that was observed because of the leakage. Bio-oil
leakage is discussed in Section 4.4.2.
Figure 4.3: Volatile Temperature Graph of Run A1
The volatile inlet temperature in the condenser was also constantly changing with
time, also shown in Figure 4.3. This meant that the mass flow rate of the volatiles
changed with time. When the mass flow rate was high the temperature drop was less
64
because of the increased heat capacity of the volatiles. It was observed in the experiment
that when the blower was turned on the volatile inlet temperature was higher than when
the blower was turned off, also shown in Figure 4.3. Turning the blower on increased the
velocity of the volatiles, thus, increased the mass flow rate. The bio-oil yield was also
observed to be high when the blower was turned on, discussed in Section 4.4.1.
There was also an instance when the volatile inlet temperature was relatively high
even when the blower was turned off. Unlike the sudden increase in temperature when
the blower was turned on, the temperature rise was gradual as shown in Figure 4.4.
Figure 4.4: Temperature Rise while Blower was Turned Off for Run A4
High volumes of bio-oil were collected during this phenomenon in the runs that it
occurred, which meant that there was an increase in the mass flow of the volatiles. Since
the blower was turned off, the only reason for the increased mass flow is that the
conditions inside reactor changed, increasing the rate of devolatilization. However, this is
beyond the scope of the present study. Based on the observations above, the volatile inlet
temperature could be considered as an indirect indication of the bio-oil yield, however,
only if the reactor temperatures remain fairly constant since bio-oil yield also depends on
the reactor temperature.[25]
High inlet temperatures had high yield and low inlet
temperatures had relatively lower yield. The volatile temperature graph with indicated
bio-oil yield is presented in Appendix F.
65
The time of some of the instances when the blower was turned on was recorded
and then located on the temperature graph of the corresponding run. Refer to Figure 4.3.
During run A1, it was recorded that the blower was turned on from 1:34-1:35. During
time 1:34-1:35 there was an abrupt rise of the inlet and exit temperatures of the volatiles,
and then a sudden drop after time 1:35 when the blower was turned off. There were
numerous accounts of this event throughout the entire experiment. It was concluded that
the abrupt rise in temperature was an indication that the blower was turned on. This
relationship between the temperature and blower was useful in the analysis of the bio-oil
yield which is discussed in Section 4.4.1.
When the blower was not turned on the exit temperature of the volatiles was
almost equal to the inlet temperature of the cooling water, considering the tolerance of
the thermocouple reading discussed in Section 4.2.2. This was observed for all the runs.
However, when the blower was turned on, there was an abrupt increase in the exit
temperature of the volatiles as shown in Figure 4.5. However, when the gas-exit-valve
was in the „slightly close‟ position while the blower was turned on, the rise in exit
temperature was not very high. There were even instances when the value of the exit
temperature was near the cooling water temperature when the blower was turned on and
the gas-exit-valve was in the „slightly close‟ position.
Figure 4.5: Volatile Exit and Cooling Water Inlet Temperatures for Run S1
66
The increase in exit temperature was mainly due to the increase in mass flow rate and can
easily be explained mathematically, as shown below. From heat balance, for an ideal
system with no heat loss, the heat released by the volatiles equals the heat absorbed by
the cooling water.
Heat Released by Volatiles = Heat Absorbed by Cooling Water
, , v v v in v ex w w wm c T T m c T
4.1
where ṁv and cv are the mass flow rate and specific heat of the volatiles, respectively; Tv,in
and Tv,ex are inlet and exit temperatures of the volatiles, respectively; ṁw and cw are the
mass flow rate and specific heat of the cooling water; ΔTw is the change in temperature of
the cooling water. Rearranging Eq. (4.1) yields the solution for the exit temperature of the
volatiles.
, ,
w w w
v ex v in
v v
m c TT T
m c 4.1
When the blower was turned on, ṁv and Tv,in are increased while cv, ṁw, cw, and ΔTw
remain fairly constant. The increase in ṁv when the blower was turned on was large
enough to significantly affect the 2nd
term in the right side of Eq. (4.1) above, thus
increasing Tv,ex. However, when the gas-exit-valve was in the „slightly close‟ position, the
increase in ṁv was not large enough to affect Tv,ex greatly. This was also true when there
was a gradual increase of the inlet temperature while the blower was turned off. In this
situation the value of Tv,ex did not seem to be affected at all.
4.2.2. Temperature of Cooling Water
As predicted in Section 3.5.3, the temperature rise of the cooling water was not
read clearly by the thermocouple. Looking at the raw data, the exit temperature was
higher than the inlet temperature, however, when the tolerance of the thermocouple is
taken into account, the inlet and exit temperatures overlap. This means that the inlet and
exit temperature could be the same, but the thermocouple recorded differently because of
67
the tolerance. Table 4.1 shows the recorded temperature and the range of the actual
values for run A2.
Table 4.1: Cooling Water Temperature Reading for Run A2
Cooling Water Inlet, °C Cooling Water Exit, °C
27.3 28.3
Range Plus (+) Minus (-) Plus (+) Minus (-)
27.9 26.6 28.9 27.6
Varying the flow rate of the cooling water from „fully open‟ to „four valve-turns‟
did not exhibit any change in the temperature. The change in temperature was observable
only when the flow was completely stopped in run S1, i.e. valve fully closed, as shown in
Figure 4.6.
Figure 4.6: Cooling Water Exit Temperature in Run S1
The exit temperature increased, although very slowly. When the valve was opened, even
just a little, the exit temperature returned to its initial value almost instantaneously. This
was because of the great difference in heat capacity rates between the two mediums, that
is,
w vC C
4.2
which is equal to
,
w w TP TP subm c m Ac
68
40.3 4,176 0.381 4.486 10 1,666.313
1,252.8 J s K 0.285 J s K
If the mass flow rate of the cooling water is decreased to 0.149 kg/s (refer to Table C.5)
and the mass flux of the volatiles in increased to 1.474 kg/m2·s (vTP = 0.851 m/s), the heat
capacity rates are
622.224 J s K 0.754 J s Kw vC C
There was also a discrepancy between the cooling water inlet temperature and the
volatile exit temperature readings. In some runs the exit temperature of the volatiles was
lower than the inlet temperature of the cooling water, which is thermodynamically
impossible with regards to the experiment set-up. Again, the reason for this discrepancy
could be the tolerance of the thermocouple reading. Table 4.2 shows the cooling water
inlet temperature and volatile exit temperature and their corresponding range of the actual
values for run S1. Another reason could be deposits of bio-oil in the thermocouple probes
which led to the inaccuracy of the volatile exit temperature reading. The effectiveness of
the condenser was not computed because of the discrepancy between the volatile exit and
cooling water inlet temperatures.
Table 4.2: Cooling Water Inlet and Volatile Exit Temperatures for Run S1
Volatile Exit, °C Cooling Water Inlet, °C
28 28.8
Range Plus (+) Minus (-) Plus (+) Minus (-)
28.6 27.4 29.4 28.1
If the thermocouples used in reading the cooling water temperature had been more
accurate and sensitive, the actual heat transferred in the condenser could have been
calculated more accurately from the cooling water since the mass flow rate and specific
heat of the cooling water can be determined accurately. The calculation of the actual heat
69
transferred discussed in Section 4.6.3 was not accurate because there were still some
assumptions made regarding the specific heat of the volatiles.
4.3. Static Pressure and Gas Velocity
Results of the static pressure measurement at the condenser inlet for runs A4 and
A5 are tabulated in Table 4.3.
Table 4.3: Inlet Static Pressure
Run
No.
Static Pressure, Pa (gage)
Gas Valve Full Open Gas Valve Slightly Close
A4 77.608 279.389
A5 38.649 186.259
The static pressures in run A5 was lower than in run A4 for both „fully open‟ and
„slightly close‟ gas-exit-valve positions because of the following reason. Runs A4 and A5
were conducted in the same day, where A4 was the first run and A5 second. The blower
was not cleaned after run A4, leaving the deposit of black viscous liquid inside the
blower casing for run A5. The black viscous liquid is discussed in Section 4.4.3. The
deposits resulted in decreased blower performance which explains the lower static
pressure measured in run A5. The deposits of black viscous liquid were observed in all
the runs. For the two gas-exit-valve positions, the static pressure was higher in the
„slightly close‟ position. This is explained below together with the gas velocities at
different gas-exit-valve positions. The pressure drop in the condenser is explained in
Section 4.6.2.
The gas velocities at gas-exit-valve „fully open‟ and „slightly close‟ were
measured for runs A4 and A5 corresponding to the static pressure measurement. The gas
velocities inside the condenser were solved as shown in Appendix D. The gas velocities
inside the condenser are tabulated below in Table 4.4. Similar to the static pressure, the
velocities in run A5, for both „fully open‟ and „slightly close‟, are lower than in run A4
because of the deposits in the blower. In contrast to the static pressure, the gas velocities
decreased when the gas-exit-valve was in the „slightly close‟ position. The reason for the
70
variation in static pressure and gas velocity with respect to gas-exit-valve position is
explained below with the aid of the fan-system curve in Figure 4.7.
Table 4.4: Gas Velocity in the Condenser Inner tube
Run
No.
Computed In-Tube Velocity, m/s
Valve Fully Open Valve Slightly Close
A4 0.851 0.325
A5 0.500 0.125
In the fan-system curve shown in Figure 4.7[12]
, the operating condition with the
gas valve in the „fully open‟ and „slightly close‟ positions are denoted by the subscripts a
and b, respectively. The point where the fan curve and the system curve (a) – gas valve
„fully open‟ – intersect is the operating point OPa, which corresponds to a static pressure
pa and velocity va. The point where the fan curve and the system curve (b) – gas valve
„slightly close‟ – intersect is the operating point OPb, which corresponds to a static
pressure pb and velocity vb. Figure 4.7 confirms the values of the static pressure and gas
velocity measurements discussed above. That is, the static pressure at OPa is lower than
OPb, and the gas velocity at OPa is higher than OPb.
Figure 4.7: Fan-System Curve[12]
71
The measured static pressures and gas velocities were used in the calculation of
the pressure drop and actual heat transfer, and the recalculation of the condenser length
which are discussed later.
4.4. Bio-oil Yield
The bio-oil collected from the marine florae feedstock was mostly composed of
the brown colored liquid with some black component at the top of the beaker as shown in
Figure 4.8. The black component was more viscous and less dense than the brown
component. More is discussed about the black component in Section 4.4.3. There were
also solid particles that were collected along with the bio-oil that settled at the bottom of
the beaker. This particles may be char particles which are carry-over from the reactor, as
discussed in Section 2.1.1.[25]
Figure 4.8: Collected Bio-oil
The average mass percentage of bio-oil based on the feedstock was 11.36%. This
is lower than the bio-oil yield from the study of Añora (2010) which was 32.73% of the
feedstock. The list of the bio-oil yield of the different feedstock in every run is tabulated
in Appendix E. The bio-oil yield in the present study was smaller than that of Añora
because the feedstock in the present study was dried in an oven prior to pyrolysis. Refer
to Esgana (2011) for the complete procedure of the pyrolysis experiment. In the
experiment of Añora the feedstock was not oven-dried prior to the experiment. Thus, the
bio-oil obtained from the experiment of Añora may contain more water content. Another
reason for the discrepancy in bio-oil yield is the pyrolysis reactor. Pyrolysis product yield
72
depend on temperature and heating rate. In Esgana‟s reactor the temperatures were varied
between different reactor zones and the heating rate was quite slow due to the large
quantities of feedstock in the reactor. Whereas in Añora‟s reactor the temperature was
well distributed because of the small amount of feedstock in the reactor and the heating
rate was not as slow.
4.4.1. Effect of Blower on Bio-oil Yield
It was observed during the experiment that there was a relative increase in the
volume of bio-oil collected in the beaker each time the blower was turned on. When the
blower was not turned on the bio-oil yield was relatively lower. The volumes of bio-oil
collected every 15 minutes for Run A7 are tabulated in Table 4.5. The bio-oil collected in
the other runs is shown in Appendix E. The data on bio-oil yield was plotted on the
volatile temperature graph of the same run.
Table 4.5: Collected Bio-oil for Run A7
Run Duration,
h:mm
Volume,
ml
Run Duration,
h:mm
Volume,
ml
1:08 Start 3:23 44
1:23 2 3:38 35
1:38 4 3:53 35
1:53 25 4:08 9
2:08 18 4:23 20
2:23 16 4:38 15
2:38 36 4:53 50
2:53 30 5:08 28
3:08 39 5:23 38
73
A portion of this graph is shown in Figure 4.9 below. It was observed that the highest
volumes collected were when the blower was frequently turned on, as indicated by the
abrupt rise in temperature in Figure 4.9. The abrupt temperature rise when the blower was
turned on was discussed in Section 4.2.1. The mass flow rate of the volatiles was
temporarily increased when the blower was turned on, thus, the bio-oil yield was also
increased for that same period. Complete volatile temperature graphs similar to Figure
4.9 are presented Appendix F.
Figure 4.9: Volatile Temperature Graph for Run A7
Also, as discussed in Section 3.4, the volatiles were not able to flow continuously
through the gas-exit-pipe of the reactor when the blower was turned off. Thus, the
volatiles containing the bio-oil did not flow to the condenser, and hence, the bio-oil was
not condensed. This is why the blower was important in the pyrolysis set-up.
74
4.4.2. Bio-oil Leakage
The undesired leakage of bio-oil in the blower meant that there was condensation
even before the volatiles entered the condenser. In run A8 there was recorded leakage in
the blower listed in Table E.8 of Appendix E. The recorded time when the leak was
observed was plotted on the volatile temperature graph of run A8, shown in Figure 4.10.
Figure 4.10: Bio-oil Leakage Plotted in Volatile Temperature Graph of Run A8
The graph showed that the volatile inlet temperature was as high as 93.4°C, as indicated
in Figure 4.10. This indicates that bio-oil starts condensation at temperatures higher than
93.4°C, which is true for the water content of bio-oil.
4.4.3. Black Viscous Liquid
The black viscous liquid was observed to stick to the blower casing and blades,
the pipe fittings, and condenser inner tube wall as shown in Figure 4.11 indicated by the
red highlights.
Figure 4.11: Unrecovered Black Viscous Liquid
a) Adapter b) Condenser c) Blower
75
This result confirms the statements about the viscous liquid that result from slow cooling,
as mentioned in Section 2.1.2. The black viscous liquid was also observed to condense in
the hopper of the reactor[13]
when the reactor was opened during loading. According to
the study of Esgana[13]
the black viscous liquid has a heating value of 26,260.19 kJ/kg,
which is very high compared to the collected brown-colored component of the oil with
heating values ranging from 103.09 kJ/kg to 628.14 kJ/kg. However, there were very
small amounts of the black viscous liquid that was observed. The black viscous liquid
that stuck in the blower was collected only when the blower was cleaned after the
experiment. Nonetheless, it seems that it is the best candidate for alternative fuel simply
because of its high heating value. Future condenser designs, therefore, must allow for
collection of the black viscous liquid.
Smaller amounts of the black viscous liquid, indicated by the red highlights in
Figure 4.12, were collected in the beaker. When the collected bio-oil was transferred to
the bottle containers for storage, some of the black viscous liquid was left in the beaker
and graduated cylinder as shown in Figure 4.13. The black viscous liquid left in the
beaker could have caused an error in weighing the bio-oil because the bio-oil was
weighed only after it was transferred to the bottle container.
Figure 4.12: Collected Black Viscous Liquid
Figure 4.13: Black Viscous Liquid Residue
76
4.5. Pyrolysis Gas
The pyrolysis gas was found to be combustible as shown in Figure 4.14. Its
combustibility was due to its methane content shown in Table 4.6 below.
Figure 4.14: Flame from Pyrolysis Gas
4.5.1. Components
The pyrolysis gas for the six types of marine florae listed below has high
concentration of carbon dioxide, and smaller concentration of methane. Results of gas
chromatograph are shown in Table 4.6. The type of gas chromatograph apparatus that
was used was limited to detect carbon dioxide and methane only.
Table 4.6: Component Percentage of Pyrolysis Gas
Marine Florae Feedstock Gas Component, %
Methane Carbon Dioxide
Seagrass w/ Binder 9.76 90.24
Pure Red Pellets 23.24 76.76
Red Raw 25.00 75.00
Green Raw 13.72 86.28
Pure Green Pellets 7.92 92.08
Brown w/ Binder 7.03 92.97
77
4.5.2. Estimate of Pyrolysis Gas Yield
The mass of the pyrolysis gas was estimated by subtracting the mass of bio-oil
and char from the original feedstock mass. That is,
mass of mass of mass of mass ofpyrolysis gas feedstock char bio oil
-
4.3
This estimate of the pyrolysis gas also includes the mass of black viscous liquid that was
not measured, as discussed in Section 4.4.3, and the residue left in the reactor. The
average mass percentage of pyrolysis gas based on the feedstock was 20.13%. This is
higher than the pyrolysis gas yield from the study of Añora (2010) which was 13.09%.
Again, the reason might be the configuration of the pyrolysis reactor, as discussed in
Section 4.4.3. The list of the pyrolysis gas yield of the different feedstock in every run is
tabulated in Table E.2 in Appendix E.
4.6. Condenser Performance
During the entire experiment the condenser seemed to be just a subcooling heat
exchanger because of undesired condensation of the bio-oil prior to the condenser inlet.
The oil leak observed in the blower was evidence that there was condensation there.
Certainly, the water content[8]
of the bio-oil was condensed in the blower. Since the bio-
oil was considered to have a condensing temperature equal to water because the
condensing temperatures of other bio-oil components were not known, only the
subcooling section of the fabricated condenser was analyzed. The desuperheating and
condensing sections were not analyzed. In spite of this, the undesired condensation in the
blower provided some clue to the condensation temperature of the bio-oil which was
discussed in Section 4.4.2.
The effectiveness of the condenser was not calculated because of the discrepancy
between the cooling water inlet temperature and volatile exit temperature, as discussed in
Section 4.2.2. Also, the actual overall heat transfer coefficient was not calculated because
of the same reason.
78
4.6.1. Condenser Material
As discussed in Section 4.4.3, there was a black viscous liquid that was observed
to adhere to the condenser walls that could cause fouling. When the condensers were
disassembled from the pyrolysis set-up to be cleaned after the experiment, less black
viscous liquid was observed in the walls of the aluminum condenser than in the stainless
condenser as shown in Figure 4.15. The task of cleaning the condensers was mainly to
remove the black viscous liquid from the condenser walls. The aluminum condenser was
easier to clean than the stainless condenser.
Figure 4.15: Comparison of Stainless and Aluminum Condensers
The inner tube of the aluminum condenser was almost entirely free of the black viscous
liquid after spraying water through the inner tube. Figure 4.16 shows the walls of the
inner tube of the aluminum condenser before and after spraying with tap water. The
stainless condenser, on the other hand, needed a test tube brush to clean the inner tube.
Figure 4.16: Aluminum Condenser
Also, there were no obvious indications of corrosion in both materials after a total
of 40.27 hours and 37.32 hours of operation for the aluminum and stainless condenser,
b) Aluminum Condenser a) Stainless Condenser
b) After spraying with water a) Before spraying with water
79
respectively. However, it was later discovered, through further literature survey, that bio-
oils can be corrosive to aluminum due to the presence of acetic acid and formic acid.[21]
Aside from the physical observations discussed above, aluminum tube is also cheaper
than stainless. Aluminum costs P70/m while stainless costs P140/m in the local market.
In terms of thermal requirements, aluminum seems to have the advantage over stainless
steel because of its higher thermal conductivity, 204 W/m·K for aluminum and 16.3
W/m·K for stainless.[15]
However, based on the recalculation discussed in Section 4.7.3,
the thermal conductivity of the condenser has little effect on the condenser length. There
was black viscous liquid that also stick to the GI pipe fittings, which was even harder to
clean. Galvanized iron is not recommended for condenser material.
4.6.2. Pressure Drop
The calculations for the pressure drop are shown in Appendix I and the values of
the pressure drop at different gas velocities are tabulated in Table I.2 to I.5. The pressure
drop in the condenser was very small, an average of 5.990 Pa for Run A4 and „full open‟
gas-exit-valve. The maximum pressure drop determined was from Run A5 and „slightly
closed‟ gas-exit-valve position, which was 6.201 Pa. The small pressure drop was not
enough to significantly affect the value of the exit density of the pyrolysis gas as
discussed in Appendix I. The change in exit density of the pyrolysis gas was 5.977 x 10-3
%. This means that the initial estimate of the two-phase density, mass flux, and heat
transfer, which is discussed in Section 4.6.3, were sufficiently accurate.
4.6.3. Actual Heat Transferred
The actual heat transferred, calculated in Appendix I.2, is much lower than the
calculated value in the initial design from Eq. (A.5) of Appendix A. In the initial design
calculation, the parameters were manipulated to yield the maximum possible heat
transfer, as discussed in Section 3.2.1. The required heat transfer calculated in Appendix
A was 3,804.734 W, and the highest value of the estimated heat transfer that occurred
during the experiment was 19.535 W, as shown in Table I.7. Refer to Appendix I.2 for
the calculations of the actual heat transferred at different gas velocities. The values of the
actual heat transferred are tabulated in Table I.7 to I.10. The amount of heat transfer is
80
directly proportional to the mass flow rate (ṁ) and the change in temperature (ΔT) of the
volatiles. The ṁ of the volatiles calculated in Appendix A was 2.3 x 10-3
kg/s while the ṁ
in Appendix I was 6.473 x 10-4
kg/s. The ΔT used in Appendix A was 79°C while the
actual ΔT was only about 30°C. The smaller than expected ΔT was due to the heat lost in
the gas-exit-pipe of the reactor and in the blower as discussed in Section 4.2.1. Both ṁ
and ΔT were higher in the initial calculation which is the reason that the heat transfer was
higher.
4.7. Results of Recalculation of the Condenser Length
4.7.1. Comparison of Initial Calculation and Recalculation
The recalculated condenser length was much longer than the initial calculation
even though the mass flow rate of the volatiles in the recalculation was lower, as
discussed in Section 4.6.3. The length of the aluminum condenser in the initial
calculation was 78.1 cm while in the recalculation, the length was 264.00 cm. The reason
for this is the small value of the convection heat transfer coefficients of the volatiles at
the desuperheating and subcooling zones. In the initial calculation the convection
coefficient of the volatiles used for the entire condenser was 5,061.875 W/m2·K. In the
recalculation, the convection coefficients at the desuperheating and subcooling zones
were 3.844 W/m2·K and 6.495 W/m
2·K, respectively, and 5,974.601 W/m
2·K at the
condensing zone. Because of the low convection coefficients in the desuperheating and
subcooling zones, the condenser must be lengthened to be able to meet the required heat
transfer duty. The relationship between the convection coefficient of the volatiles and the
cooling water is discussed in Section 4.7.4. The zone with the longest length is the
subcooling zone because of the large required ΔT (100°C to 31°C). The length of the
subcooling zone is about 94 % of the entire condenser length. Refer to Table H.8 in
Appendix H.
The experiment results, however, were not comparable with the results of the
recalculation because the designed operating temperatures were not realized. The inlet
temperature of the volatiles in the experiment was much lower than expected, as
discussed in Section 4.2.1.
81
4.7.2. Effect of Flow Velocity
If the flow velocity is increased, while keeping all other variables constant, the
required condenser length and the pressure drop are increased as shown in Figure 4.17.
Small variations in flow velocity have large effect on the required condenser length. The
flow velocity seems to be the main variable affecting the required length of the
condenser. Therefore, proper blower and piping design is necessary to ensure that
condensing system will be able to manage the variations in flow velocity.
Figure 4.17: Flow Velocity, Condenser Length, Pressure Drop
4.7.3. Effect of Thermal Conductivity of Condenser Tube
The result of the recalculation was different from the initial calculation where the
thermal conductivity of the tube material had a significant effect on the condenser length.
The thermal conductivity of the tube has negligible effect on the length. If the tube
material is 25% Cr & 20% Ni which has a thermal conductivity of 12.80 W/m·K the
resulting length is 264.26 cm; if the material is pure silver which has a thermal
conductivity of 419 W/m·K the length is 263.99 cm. This is equivalent to a 0.10% change
in length. In terms of thermal conductivity, the two materials mentioned are the two
extremes shown in Table A-2 of reference [15]. Therefore, there is little significance in
using a material with high thermal conductivity.
0
20
40
60
80
100
120
0
200
400
600
800
1,000
1,200
1,400
0 0.25 0.5 0.75 1
Co
nden
ser
Len
gth
, cm
Flow Velocity, m/s
Length
Pressure
Drop
Pre
ssure
Dro
p, P
a
82
4.7.4. Effect of Cooling Water
High convection heat transfer coefficients of the cooling water have no significant
effect on the required heat transfer length as shown in Figure 4.18. Since the overall heat
transfer coefficient is governed mostly by fluid with the lower convection coefficient,
which is the volatiles, increasing the cooling water convection coefficient to substantially
high values has little effect on the overall heat transfer coefficient.
Figure 4.18: Cooling Water Convection Coefficient and Condenser Length
The overall heat transfer coefficient U can be calculated from Eq. (4.4).
1
ln1 1
2
o i
i v o w
U =d d
+ +Ah πkL A h
4.4
where hw represents the convection coefficient of the cooling water. For the sake of
illustration, let the 1st and 2
nd terms of the denominator, and Ao assume values of unity. If
hw = 1,000,
-3
1= 0.50
1+1+1×10U
0
1
2
3
4
5
6
260 265 270 275
Co
nvec
tio
n C
oef
ficie
nt,
10
3W
/m2·K
Condenser Length, cm
83
That is, for this example only,
lim 10.5
ln1 1
2
o iw
i v o w
Ud dh
Ah πkL A h
+ +
For smaller values of hw, however, the overall heat transfer coefficient U decreases. For
example, if hw = 10,
10.476
1 1 0.1U =
+ +
Therefore, if the convection heat transfer coefficient of the volatiles is held
constant, increasing the convection coefficient of the cooling water to substantially high
values do not significantly affect the overall heat transfer coefficient and the condenser
length. Rather, the overall heat transfer coefficient will reach an upper limit. However,
decreasing the convection coefficient of the cooling water below some critical value can
significantly affect the overall heat transfer coefficient and the condenser length.
84
CHAPTER 5
CONCLUSIONS AND RECOMMENDATIONS
5.1. Conclusions
The fabricated condensers were able to perform their designed purpose that was to
extract the bio-oil from the volatiles. The bio-oil was successfully collected and the rate
of bio-oil yield was determined. The condensers were effective in cooling the volatiles to
a temperature that was almost equal to the temperature of the cooling water most of the
time. The actual numerical value of the effectiveness was not determined because of
discrepancies in temperature reading between the cooling water inlet and volatile exit
temperatures. With regards to condenser material, the aluminum tube was better than the
stainless steel tube in terms of cleanability.
The installed centrifugal blower was a vital equipment of the pyrolysis system. It
was able to direct the flow of the volatiles to the condenser instead of getting trapped
inside the reactor and then escaping to the atmosphere. However, the blower provided an
extra heat transfer area and undesired condensation was observed to have taken place in
the blower.
The rate of bio-oil yield collected in the beaker was found to be fluctuating and
partially dependent on the flow velocity. Small amounts of the black viscous liquid were
also collected in the beaker. This black viscous liquid was also observed to adhere to the
walls of the condenser and blower which made its collection and measurement difficult.
The condensation temperature of the bio-oil was observed to be higher than 93.4°C
which is certainly true for its water content. The mass percentages of the bio-oil and
pyrolysis gas yield did not match the results of Añora (2010) because of the dissimilarity
in process conditions. The components of the pyrolysis gas that were detected are Carbon
Dioxide (CO2) and Methane (CH4).
The revised solution for the condenser produced the following results: The effect
of the convection heat transfer coefficient of the cooling water was negligible at values
85
greater than 3,000 W/m2·K. The thermal conductivity of the condenser tube had
insignificant effect on the required condenser length. The required length of the
condenser was mostly dependent on the flow velocity of the volatiles. The result of the
recalculation showed that required length was impractical for a double-pipe condenser.
5.2. Recommendations
Suggestion for improvements of the present pyrolysis set-up and future studies
that will contribute to the improvement of the condenser design are discussed below:
1) For future research, better and more accurate measuring instruments are
recommended. Thermocouples should be able to read small variations in temperature,
especially those that are used for reading the cooling water temperature. The actual heat
transfer would be estimated more accurately if it is calculated based on the heat absorbed
by the cooling water. Digital pressure sensors with datalogging are recommended.
Simultaneous measurement of the inlet and exit static pressures should be done. A digital
velocity meter that is capable of measuring a wide range of flow velocities with good
accuracy is also suggested.
2) The gas-exit-pipe of the reactor should be revised to prevent premature condensation
and to allow the volatiles to flow more freely out the reactor. The diameter of the gas-
exit-pipe can be increased so that the flow is not obstructed. Bridgwater (1999) suggests
maintaining the transfer lines from the reactor to the condenser at a high enough
temperature to minimize oil deposition. The temperature of the volatiles while still in the
blower should be high enough to avoid condensation and deposition of oil in the blower
which reduces the performance of the blower. Another way to prevent oil deposition in
the blower is to place the blower after the condenser. Also, the pyrolysis system should
include filters to capture solid particles contained in the volatiles which are carry-over
from the reactor.
3) The thermophysical properties of the bio-oil should be determined in order to be able
to design the condenser with more accuracy. The actual condensing temperature of the
86
bio-oil should also be determined. The acidity of the bio-oil should also be determined to
be able properly select the condenser material. Literatures suggest that the pyrolysis gas
has more components besides CO2 and CH4. The other gas components should be
determined to be able to design the condenser more accurately.
4) The blower should be selected properly to match the operating conditions of the
pyrolysis system. The capacity of the blower should coincide with the rate of
devolatilization. That is, the rate at which the blower sucks out the volatiles from the
reactor should be the same as the rate at which the volatiles are liberated from the
feedstock. Matching the suction rate to the devolatilization rate ensures that there is
continuous motion of the volatiles out the reactor, reducing the residence time of the
volatiles. Less residence time inside the reactor results to more liquid yield.[7]
The blower
should also be able to provide enough pressure to drive the volatiles through and out the
condenser.
5) The results of the recalculation show that at high flow velocities, the condenser length
requirement is very long and the use of a double-pipe condenser would be impractical. In
this case the Shell-and-Tube condenser is recommended. The design calculations
presented in this study can be adapted to shell-and-tube because of its tubular geometry.
In the same manner, the volatiles would flow in the tube side and the cooling water in the
shell side. The tubes can also be tilted to enhance condensation. Moreover, the tube side
can be cleaned mechanically. The volatile flow has to be metered so that the mass is
distributed properly among the several tubes. Another type of condenser that seems
appealing is the Gasketed Plate or Plate and Frame heat exchanger. This type of heat
exchanger has high heat transfer coefficients and high local shear which minimizes
fouling. Its characteristics make it well suited for heat exchange duties involving
pyrolysis volatiles, which have low convection coefficient (due to presence of
noncondensable gases) and high potential of fouling.
6) The thermal design calculation can be improved further by using a Separated Two-
Phase model instead of the Homogeneous Two-Phase model. The assumption of the
87
homogeneous two-phase model that the flow velocities of the bio-oil and pyrolysis gas
are equal was inaccurate, especially in the subcooling zone. The viscosities of the oil and
gas differ by a large value in the subcooling zone. Utilization of the separated flow
model, however, requires data on the actual flow velocity of the liquid component (bio-
oil), which was not determined in this study. The flow regime in the desuperheating,
condensing, and subcooling zones must also be known. Also, the bio-oil was composed
of two immiscible liquid components, which are the brown-colored component and the
black viscous component. The reader is referred to Bird (2002)[5]
for flow of two adjacent
immiscible liquids.
88
Appendices
Appendix A. Calculation of Initial Condenser Design
A.1. Required Heat Transfer
The solution presented here is for aluminum condenser only. The same solution was used to calculate for the size of the
stainless condenser. The mass flow rate of the volatiles was estimated based on the experiment results of Añora (2010). Only the
solution for Pure Brown Pellets is shown here because it had the highest value of heat rejection. The mass flow rates were estimated
using Eq. (A.1) and (A.2). Values for %bo, %pg, and t are shown in Appendix E; m is 5 kg.
%bo
bo mm
t
A.1
where: %bo = 0.36
t = 23 sec
-30.36 5 kg1.304×10 kg s
23 secbom
%pg
pg mm
t
A.2
where: %pg = 0.11
t = 23 sec
-40.11 5 kg= 3.99×10 kg s
23 secpgm
The heat rejection of the bio-oil was computed from Eq. (A.3). Since the bio-ol was assumed to have properties equivalent to water,
the values for hi and hsat were obtained from steam tables at 110°C and 100°C, respectively, and at atmospheric pressure; hL and cw are
the latent heat of vaporization and specific heat of water, respectively.
bo bo i sat fg w sat v,exQ = m h - h +h +c T T
A.3
where: ṁbo = 1.304x10-3 kg/s
hi = 2,696,200 J/kg
hsat = 2,676,100 J/kg
hfg = 2,257,000 J/kg
cw = 4,195 J/kg·K
Tsat = 100°C
Tv,ex = 31°C
-3= 1.304×10 2,696,200 2,676,100 + 2,257,000 + 4,195 100 31boQ
= 3,347.680 WboQ
The pyrolysis gas was assumed to be proportions of carbon dioxide, carbon monoxide, hydrogen, and methane. The percentages of
each gas component were determined by trial and error. It was determined from Eq. (A.4) that the maximum possible amount of heat
that must be rejected occurred when the pyrolysis gas was composed solely of hydrogen gas.
pg pg p v,in v,exQ = ym c T T
A.4
2 2pg H pg p,H v,in v,exQ = y m c T T
where: ṁpg = 3.99x10-4 kg/s
89
y = 1
2p,Hc = 14,500 J/kg·K
Tv,in = 110°C
Tv,ex = 31°C
-4= 1 3.99×10 14,500 110 31pgQ
= 457.054 WpgQ
The total required heat transfer in the condenser is the sum of Qbo and Qpg, as shown in Eq. (A.5).
bo pgQ= Q +Q
A.5
= 3,347.680 W + 457.054 WQ
= 3,804.734 WQ
A.2. Logarithmic Mean Temperature Difference
At the cooling water exit side to the lower reservoir, the exit velocity of the cooling water was solved from Eq. (A.6). The
elevation head z1 from the upper reservoir to the lower reservoir was estimated to be 1.2 m.
2 12v = gz
A.6
2 = 2 9.81 1.2v
2 = 4.852 m sv
The volume flow rate of the cooling water, which is used in Eq. (A.8), was computed from Eq. (A.7) below, where d2,i is equal to
0.2096 m and v2 was solved above. Refer to Appendix B for the dimensions of the condenser tubes.
2
2 24
,i
πV = v d
A.7
2
= 3.641 0.020964
πV
-3 3=1.67×10 m sV
In order to compute the mean temperature difference in the condenser, the exit temperature of the cooling water was determined from
Eq. (A.8).
w,ex w,in
w w
QT = +T
ρ Vc
A.8
where: Q = 3,804.734 W
ρw = 995.26 kg/m3
cw = 4,176 J/kg·K
Tw,in = 30°C
-3
3,804.734= + 30
995.26 1.67×10 4,176w,exT
= 30.546 °Cw,exT
The mean temperature difference is computed below in Eq. (A.9).
1 2
1
2
Δ ΔΔ
Δln ln
Δ
v,in w,ex v,ex w,in
lm
v,in w,ex
v,ex w,in
T T T TT TT = =
T T T
T T T
A.9
where: Tv,in = 110°C
90
Tv,ex = 31°C
Tw,in = 30°C
Tw,ex = 30.546°C
110 30.546 31 30Δ =
110 30.546ln
31 30
lmT
Δ =17.931°ClmT
A.3. Convection Heat Transfer Coefficient
From the continuity equation, which is simplified in Eq. (A.10), the velocity of the cooling water in the annular space
between the condenser inner tube and outer tube was solved.
2
2 22 2
2 2= = ,i
w
annulus i o
v dv Av
A D d
A.10
where: d2,i = 0.02096 m
Di = 0.03508 m
Do = 0.0254 m
v2 = 4.842 m/s
2
2 2
4.852 0.02096=
0.03508 0.0254wv
= 3.641m swv
The hydraulic diameter of the annulus was solved from Eq. (A.11). The values of Di and do were given above.
2 2
i oH
o
D - dD =
d
A.11
2 2
0.03508 0.0254=
0.0254HD
= 0.023 mHD
For annular flow, Re < 10,000 is considered as turbulent flow. The flow of the cooling water was found to be turbulent as shown
below in Eq. (A.12).
Re w w Hw
w
ρ v D=
μ
A.12
where: ρw = 995.26 kg/m3
vw = 3.641 m/s
DH = 0.023 m
μw = 8.03x10-4 kg/m·s
-4
995.26 3.641 0.023Re =
8.03×10w
Re =104,018.35w
For turbulent flow, the Nusselt number and convection heat transfer coefficient are solved from Eq. (A.13) and (A.14), respectively.
Rew was solved above and Prw is the Prandtl number of water at 30°C.
0.8 0.4Nu = 0.023 Re Prw w w
A.13
91
0.8 0.4
Nu = 0.023 104,018.35 5.412w
Nu = 539.671w
Nuw ww
H
kh =
D
A.14
where: kw = 0.619 W/m·°C
539.671 0.619=
0.023wh
2=12,533.625 W m Kwh
Eq. (A.16) was used for solving the convection heat transfer coefficient of the volatiles during condensation is valid for Re < 35,000
only. The Reynolds number was solved from Eq. (A.15) and was found to be less than 35,000, hence Eq. (A.16) is valid. The absolute
viscosity μv was assumed to be equal to water at 100°C.
4Re
bo pgi vv
v i v
m +md m= =
Aμ πd μ
A.15
where: di = 0.0239 m
ṁbo + ṁpg = 2.3x10-3 kg/s
μv = 2.82x10-4 kg/m·s
-3
-4
4 2.3×10Re =
0.0239 2.82×10v
π
Re = 434.501v
1/4 1/4
0.68sin0.555
3
fg w sat iw w v w
v
w i sat i
h + c T Tρ ρ ρ g α kh =
μ d T T
A.16
Eq. (A.16) cannot be solved directly because of the unknown inside wall temperature Ti. The solution for hv requires two equations.
For steady state heat transfer
v,in i i o o w
v t w
T T T T T TQ= = =
R R R
A.17
where:
1v
i v
R =πd Lh
A.18
ln
2
o i
t
t
d dR =
πk L
A.19
1w
o w
R =πd Lh
A.20
The 1st equation is obtained as follows.
v,in i i o
v t
T T T T=
R R
1/4
3 ln0.68sin0.555
2
1/ 4
v,in i i o ifg w sat iw w v w
i o
w i sat i t
T T d d dh + c T Tρ ρ ρ g α kT = +T
μ d T T k
A.21
92
The 2nd equation is obtained similarly.
i o o w
t w
T T T T=
R R
2
ln
2
ln
t io w w
o i
ot
o w
o i
k Td h T +
d dT =
kd h +
d d
A.22
Substituting the 2nd Equation to the 1st Equation yields
1/4 1/43
2
ln0.68sin ln0.555
22
ln
t io w w
v,in i i o ifg w sat iw w v w o i
itw i sat i t
o w
o i
k Td h T +
T T d d dh + c T Tρ ρ ρ g α k d dT = +
kμ d T T kd h +
d d
A.23
The numerical values of the variables in Eq. (A.23) are shown in Table A.1. Tube diameters are shown in Appendix B.
Table A.1: Values of Variables in Eq. (A.23)
Variable Numerical Value Unit Variable Numerical Value Unit
ρw 958.1 kg/m3 hfg 2,257,000 J/kg
ρv 0.5506 kg/m3 hw 12,533.625 W/m2-K
μw 2.825x10-4 kg/m-s Tw 30 °C
α 20 deg Tsat 100 °C
kw 0.6816 W/m-K Tv,in 110 °C
kt 204 W/m-K g 9.81 m/s2
cw 4,195 J/kg-K
Applying Bisection Method to Eq. (A.23) to solve for Ti yields
= 49.931°CiT
Ti is substituted back to Eq. (A.16) to solve for the volatile convection heat transfer coefficient. 2= 5,061.875 W m Kvh
A.4. Condenser Length
The computed values of Q, ΔTlm, hw, and hv were substituted to Eq. (A.24), which solves the length of the aluminum
condenser.
ln1 1
Δ 2
o i
lm i v t o w
d dQL = + +
π T d h k d h
A.24
where: Q = 3,804.734 W
ΔTlm = 17.931°C
hw = 12,533.625 W/m2·°C
hv = 5,061.875 W/m2·°C
kt = 204 W/m·°C
do = 0.0254 m
di = 0.0239 m
ln 0.0254 0.02393,804.734 1 1= + +
17.931 0.0239 5,061.875 2 204 0.0254 12,533.625L
π
= 0.781m = 78.1cmL
The same procedures above were followed in calculating the length of the stainless steel condenser. The length of the
stainless condenser was determined to be 0.999 m or 99.9 cm.
93
Appendix B. Fabricated Condenser Parts and Assembly
B.1. Dimensions and Parts
Figure B.1: Aluminum Condenser
Figure B.2: Stainless Condenser
Note: All dimensions above are in cm
Table B.1: List of Parts
Part # Part Name Material Size
1 Outer Tube GI pipe 1-1/4 in.
2 Inner Tube Aluminum/Stainless pipe 1 in.
3 Cross Tee GI cross tee 1-1/4 in.
4 Adapter Nipple GI nipple 1 x 1 in.
5 Water Inlet/Exit Nipples GI nipple ¾ x 2 in.
6 Bushing Reducer GI bushing reducer 1-1/4 x 1 in.
94
Figure B.3: Exploded View of Condenser
Table B.2: Inner-Tube Actual Dimensions
Aluminum (1 in.) Stainless (1 in.)
Inner Dia. (di), m Outer Dia. (do), m Inner Dia. (Di), m Outer Dia. (Do), m
0.0239 0.0254 0.0226 0.0254
Table B.3: Outer-Tube Actual Dimensions
Outer tube (1-1/4 in.) Water Inlet/Exit Tube (3/4 in.)
Inner Dia. (Di), m Outer Dia. (Do), m Inner Dia. (d2,i), m
0.03508 0.0425 0.02096
B.2. Condenser Accessories
The adapter is an accessory of the condenser where the bio-oil and pyrolysis gas are isolated. Figure B.4 shows the adapter.
By density difference the heavier liquid bio-oil flows below the lighter pyrolysis gas. When they reach the adapter, the bio-oil drops
down, as indicated in Figure B.4, to the beaker. When the oil-valve is closed the pyrolysis gas flow directly out through the gas-exit-
valve. The inner diameter of the gas-exit-valve is 7.5 mm.
Figure B.4: Adapter
95
The static pressure tap was another accessory of the condenser which was used in measuring the static pressure of the flow.
The manometer tube was attached to the valve indicated in Figure B.5.
Figure B.5: Static Pressure Tap
B.3. Thermocouple Probes and Pressure Taps
Figure B.6: Position of Thermocouple Probes
Figure B.7: Position of Static Pressure Taps
a) Aluminum Condenser
b) Stainless Condenser
96
B.4. Condenser Tilt Angle
The actual tilt angle was measure by using a water-hose level gage similar to Figure 3.13 and a straight edge, in this case, a
triangle. The two points in the water-hose level gage where the top of the water column rest represent two points on the horizontal.
These two points were connected by the triangle. The triangle then represents a horizontal line. A photograph of this set-up, shown in
Figure B.8, was taken with the camera positioned at approximately the same elevation as the set-up. In the photograph, lines parallel
to the triangle and the condenser were drawn. The angle between these lines, which was measured with a protractor, is the tilt angle of
the condenser.
Figure B.8: Condenser Tilt Angle
97
Appendix C. Cooling Water Flow Rate Measurements
Table C.1: Mass Flow Rate for Fully Open
Trial Mass, kg Time, sec Mass Flow, kg/s
1 5.734 11.74 0.488
2 5.432 11.37 0.478
3 4.984 10.4 0.479
4 5.374 11.64 0.462
5 5.058 11.63 0.435
Average 0.468
Table C.2: Mass Flow Rate for One Valve-Turn
Trial Mass, kg Time, sec Mass Flow, kg/s
1 5.364 11.4 0.471
2 4.936 10.58 0.467
3 4.954 10.13 0.489
4 4.842 10.3 0.470
5 4.852 10.33 0.470
Average 0.473
Table C.3: Mass Flow Rate for Two Valve-Turns
Trial Mass, kg Time, sec Mass Flow, kg/s
1 4.252 10.43 0.408
2 4.166 10.35 0.403
3 4.2 10.26 0.409
4 4.26 10.35 0.412
5 4.238 10.41 0.407
Average 0.408
Table C.4: Mass Flow Rate for Three Valve-Turns
Trial Mass, kg Time, sec Mass Flow, kg/s
1 2.99 10.06 0.297
2 3.146 10.46 0.301
3 3.094 10.29 0.301
4 3.136 10.37 0.302
5 3.072 10.35 0.297
Average 0.300
Table C.5: Mass Flow Rate for Four Valve-Turns
Trial Mass, kg Time, sec Mass Flow, kg/s
1 1.59 10.35 0.154
2 1.48 10.17 0.146
3 1.618 10.92 0.148
4 1.524 10.36 0.147
5 1.538 10.25 0.150
Average 0.149
98
Appendix D. Static Pressure and Gas Velocity Measurements
D.1. Static Pressure Measurements
Table D.1: Manometer Reading for Run A4 (Gas-Exit-Valve Full Open)
Manometer Tube Manometer Reading, inches H2O
Average, inches H2O Pt. 1 Pt. 2
Upper 1-1/2 1-14/16 1.6875
Lower 2-2/16 2-1/2 2.3125
Table D.2: Manometer Reading for Run A4 (Gas-Exit-Valve Slightly Close)
Manometer Tube Manometer Reading, inches H2O
Average, inches H2O Pt. 1 Pt. 2
Upper 10/16 1 0.8125
Lower 2-14/16 3-4/16 3.0625
Table D.3: Manometer Reading for Run A5 (Gas-Exit-Valve Full Open)
Manometer Tube Manometer Reading, inches H2O
Average, inches H2O Pt. 1 Pt. 2
Upper 1-15/16 2-5/16 2.12375
Lower 1-10/16 2 1.8125
Table D.4: Manometer Reading for Run A5 (Gas-Exit-Valve Slightly Close)
Manometer Tube Manometer Reading, inches H2O
Average, inches H2O Pt. 1 Pt. 2
Upper 1 1-6/16 1.1875
Lower 2-8/16 2-14/16 2.6875
Since the manometer was inclined at approximately 30° from the horizontal, Eq. (D.1) was used to solve the actual height
of the water column.
Δ sin 30°t th u l
D.1
Δ 1.6875 2.3125 sin 30°h
2Δ 0.3125 in. H Oh
Converting the height of the water column to,
2
101325 PaΔ
408 in.H Osp h
2
2
101325 Pa0.3125 in. H O
408 in.H Osp
77.608 Pasp
99
D.2. Gas Velocity Measurements
Table D.5: Gas Exit Velocity
Run No. Gas Exit Velocity, fpm
Fully Open Slightly Close
A4 1,700 650
A5 1,000 250
i i nozzle exAv A v
D.2
2
2
nozzle ex nozzle exi
i i
A v d vv
A d
2-3
2
7.5×10 m 1,700 ft min 1m 1min
3.28 ft 60 sec0.0239 miv
0.851m siv
Table D.6: Velocity Inside Inner-tube
Run No. Gas Exit Velocity, m/s
Fully Open Slightly Close
A4 0.851 0.325
A5 0.500 0.125
100
Appendix E. Pyrolysis Products
E.1. Periodic Bio-oil Volume Measurement
Table E.1: Bio-oil Volume Collected for Run A1
Run Duration,
h:mm Volume, ml
Run Duration,
h:mm Volume, ml
0:42 Start 3:12 55
0:57 34 3:27 42
1:12 33 3:42 30
1:27 31 3:57 41
1:42 26 4:12 49
1:57 12 4:27 47
2:12 20 4:42 36
2:27 16 4:57 32
2:42 42 5:12 28
2:57 44
Total: 618 ml
Table E.2: Bio-oil Volume Collected for Run A2
Run Duration,
h:mm Volume, ml
Run Duration,
h:mm Volume, ml
0:10 Sart 2:25 34
0:25 5 2:40 22
0:40 3 2:55 35
0:55 5 3:10 35
1:10 8 3:25 32
1:25 15 3:40 34
1:40 14 3:55 48
1:55 38 4:10 16
2:10 23
Total: 367 ml
Table E.3: Bio-oil Volume Collected for Run A3
Run Duration,
h:mm Volume, ml
Run Duration,
h:mm Volume, ml
0:07 Start 1:22 17
0:22 26 1:37 60
0:37 20 1:52 22
0:52 17 2:07 36
1:07 18
Table E.4: Bio-oil Volume Collected for Run A4
Run Duration,
h:mm Volume, ml
Run Duration,
h:mm Volume, ml
1:39 Start 2:39 70
1:54 29 2:54 30
2:09 39 3:09 35
2:24 33 3:24 14
Leak: 4 ml
Total: 254 ml
Table E.5: Bio-oil Volume Collected for Run A5
Run Duration,
h:mm Volume, ml
Run Duration,
h:mm Volume, ml
0:11 Start 3:11 12
0:26 31 3:26 63
0:41 33 3:41 38
0:56 21 3:56 36
1:11 12 4:11 21
1:26 12 4:26 13
1:41 16 4:41 18
1:56 28 4:56 21
2:11 13 5:11 27
2:26 20 5:26 21
2:41 33 5:41 39
2:56 13 5:56 38
Total: 579 ml
Table E.6: Bio-oil Volume Collected for Run A6
Run Duration,
h:mm Volume, ml
Run Duration,
h:mm Volume, ml
1:56 0 4:26 26
2:11 15 4:41 44
2:26 25 4:56 27
2:41 24 5:11 26
2:56 28 5:26 31
3:11 37 5:41 24
3:26 45 5:56 32
3:41 35 6:11 30
3:56 42 6:26 29
4:11 28 6:41 24
Leak: 17 ml
Total: 589 ml
101
Table E.7: Bio-oil Volume Collected for Run A7
Run Duration,
h:mm Volume, ml
Run Duration,
h:mm Volume, ml
1:08 Start 3:23 44
1:23 2 3:38 35
1:38 4 3:53 35
1:53 25 4:08 9
2:08 18 4:23 20
2:23 16 4:38 15
2:38 36 4:53 50
2:53 30 5:08 28
3:08 39 5:23 38
Leak: 24 ml
Total: 468 ml
Table E.8: Bio-oil Volume Collected for Run A8
Run Duration, h:mm Volume, ml Leakage, ml
0:20 Start
0:35 24 52
0:50 29 51
1:05 11 21
1:20 29 25
1:35 6 18
1:50 24 23
2:05 11 10
2:20 58 18
2:35 41 9
2:50 28 4
3:05 18 4
3:20 40 4
3:35 26 6
3:50 10
4:05 13
4:20 32
Leak: 245 ml
Total: 645 ml
Table E.9: Bio-oil Volume Collected for Run S5
Run Duration,
h:mm Volume, ml
Run Duration,
h:mm Volume, ml
0:53 Start 3:38 49
1:08 28 3:53 41
1:23 29 4:08 18
1:38 29 4:23 26
1:53 29 4:38 23
2:08 33 4:53 55
2:23 25 5:08 57
2:38 24 5:23 35
2:53 24 5:38 34
3:08 15 5:48 28
3:23 23
Total: 625 ml
Table E.10: Bio-oil Volume Collected for Run S6
Run Duration,
h:mm Volume, ml
Run Duration,
h:mm Volume, ml
0:08 Start 3:08 72
0:23 75 3:23 39
0:38 29 3:38 23
0:53 18 3:53 16
1:08 13 4:08 58
1:23 16 4:23 46
1:38 18 4:38 30
1:53 10 4:53 34
2:08 16 5:08 39
2:23 13 5:23 58
2:38 8 5:38 58
2:53 12 5:53 55
Total: 756 ml
Table E.11: Bio-oil Volume Collected for Run S7
Run Duration,
h:mm Volume, ml
Run Duration,
h:mm Volume, ml
2:17 Start 3:47 66
2:32 11 4:02 51
2:47 11 4:17 40
3:02 28 4:32 45
3:17 27 4:43 12
3:32 31
Total: 322 ml
102
E.2. Bio-oil and Pyrolysis Gas Yield
Table E.12: Bio-oil Volume Collected for Run S8
Run Duration,
h:mm Volume, ml
Run Duration,
h:mm Volume, ml
0:06 Start 2:36 10
0:21 46 2:51 19
0:36 36 3:06 15
0:51 22 3:21 40
1:06 40 3:36 44
1:21 27 3:51 52
1:36 23 4:06 48
1:51 13 4:21 30
2:06 12 4:36 48
2:21 25
Total: 550 ml
Table E.13: Mass of Marine Florae Feedstock and
Pyrolysis Products
Run # Type of Feedstock Weight, kg
Feedstock Char Oil Gas
A1 Pure Green Pellets 5.000 2.172 0.600 2.228
A2 Pure Red Pellet 4.474 3.639 0.346 0.489
A4 Pure Green Pellets 5.000 3.768 0.228 1.004
A5 Red Raw 5.000 3.334 0.564 1.102
A6 Pure Seagrass Pellets 4.968 3.174 0.588 1.206
A7 Pure Brown Pellets 5.000 3.700 0.452 0.848
A8 Seagrass with Binder 5.000 3.406 0.638 0.956
S1 Brown with Binder 5.000 3.23 0.652 1.118
S2 Seagrass with Binder 5.000 2.546 0.984 1.470
S3 Red Raw 3.918 2.954 0.408 0.556
S4 Pure Red Pellets 3.776 2.949 0.554 0.273
S5 Pure Green Pellets 5.095 3.984 0.606 0.505
S6 Green Raw 4.998 3.296 0.724 0.978
S7 Pure Brown Pellets 4.858 3.570 0.310 0.978
S8 Pure Seagrass Pellets 5.000 3.410 0.530 1.060
Table E.14: Mass Percentage of Pyrolysis Products
Run # Type of Feedstock Product Percentage, %
Char Oil Gas
A1 Pure Green Pellets 43.44 12.00 44.56
A2 Pure Red Pellet 81.34 7.73 10.93
A4 Pure Green Pellets 75.36 4.56 20.08
A5 Red Raw 66.68 11.28 22.04
A6 Pure Seagrass Pellets 63.89 11.84 24.28
A7 Pure Brown Pellets 74.00 9.04 16.96
A8 Seagrass with Binder 68.12 12.76 19.12
S1 Brown with Binder 64.60 13.04 22.36
S2 Seagrass with Binder 50.92 19.68 29.40
S3 Red Raw 75.40 10.41 14.19
S4 Pure Red Pellets 78.10 14.67 7.23
S5 Pure Green Pellets 78.19 11.89 9.91
S6 Green Raw 65.95 14.49 19.57
S7 Pure Brown Pellets 73.49 6.38 20.13
S8 Pure Seagrass Pellets 68.20 10.60 21.20
Average 68.51 11.36 20.13
103
Table E.15: Density of Bio-oil
Run # Type of Feedstock Density, kg/m3
A1 Pure Green Pellets 970.87
A2 Pure Red Pellet 942.78
A4 Pure Green Pellets 897.64
A5 Red Raw 974.09
A6 Pure Seagrass Pellets 998.30
A7 Pure Brown Pellets 965.81
A8 Seagrass with Binder 989.15
S1 Brown with Binder 1018.75
S2 Seagrass with Binder 1004.08
S3 Red Raw 1020.00
S4 Pure Red Pellets 1045.28
S5 Pure Green Pellets 969.60
S6 Green Raw 957.67
S7 Pure Brown Pellets 962.73
S8 Pure Seagrass Pellets 963.64
Average 978.69
E.3. Product Composition and Residence Time from Añora (2010)
Table E.16: Mass Percentage and Residence Time for Green Algae[1]
Green 80/20 Pellets Pure Green Pellets
Trial 1st 2nd 3rd 1st 2nd 3rd
%bo 0.32 0.32 0.32 0.4 0.4 0.4
%pg 0.16 0.16 0.16 0.08 0.08 0.08
t, min 33 32 30 28 27 29
Table E.17: Mass Percentage and Residence Time for Red Algae[1]
Red 80/20 Pellets Pure Red Pellets
Trial 1st 2nd 3rd 1st 2nd 3rd
%bo 0.15 0.15 0.15 0.24 0.24 0.24
%pg 0.32 0.32 0.32 0.21 0.21 0.21
t, min 17 23 25 26 23 25
Table E.18: Mass Percentage and Residence Time for Brown Algae[1]
Brown 80/20 Pellets Pure Brown Pellets
Trial 1st 2nd 3rd 1st 2nd 3rd
%bo 0.36 0.36 0.36 0.36 0.36 0.36
%pg 0.11 0.11 0.11 0.11 0.11 0.11
t, min 29 25 26 55 23 28
Table E.19: Mass Percentage and Residence Time for Seagrass[1]
Seagrass 70/30 Pellets Pure Seagrass Pellets
Trial 1st 2nd 3rd 1st 2nd 3rd
%bo 0.32 0.32 0.32 0.28 0.28 0.28
%pg 0.19 0.19 0.19 0.23 0.23 0.23
t, min 61 28 33 47 29 30
104
Appendix F. Volatile Temperature Graph
F.1. Volatile Temperature Graph with Plotted Periodic Bio-oil Yield
Fig
ure
F.1
: V
ola
tile
Tem
per
ature
Gra
ph o
f R
un A
1
105
Fig
ure
F.2
: V
ola
tile
Tem
per
ature
Gra
ph o
f R
un A
2
106
Fig
ure
F.3
: V
ola
tile
Tem
per
ature
Gra
ph o
f R
un A
3
Fig
ure
F.4
: V
ola
tile
Tem
per
ature
Gra
ph o
f R
un A
4
107
Fig
ure
F.5
: V
ola
tile
Tem
per
ature
Gra
ph o
f R
un A
5
108
Fig
ure
F.6
: V
ola
tile
Tem
per
ature
Gra
ph o
f R
un A
6
109
Fig
ure
A.7
: V
ola
tile
Tem
per
ature
Gra
ph o
f R
un A
7
110
Fig
ure
F.8
: V
ola
tile
Tem
per
ature
Gra
ph o
f R
un A
8
111
Fig
ure
F.9
: V
ola
tile
Tem
per
ature
Gra
ph o
f R
un S
5
112
Fig
ure
F.1
0:
Vola
tile
Tem
per
ature
Gra
ph o
f R
un S
6
113
Fig
ure
F.1
1:
Vola
tile
Tem
per
ature
Gra
ph o
f R
un S
7
114
Fig
ure
F.1
2:
Vola
tile
Tem
per
ature
Gra
ph o
f R
un S
8
115
F.2. Volatile Temperature Graph without Plotted Periodic Bio-oil Yield
Fig
ure
F.1
3:
Vola
tile
Tem
per
ature
Gra
ph o
f R
un S
1
116
Fig
ure
F.1
4:
Vola
tile
Tem
per
ature
Gra
ph o
f R
un S
2
117
Fig
ure
F.1
5:
Vola
tile
Tem
per
ature
Gra
ph o
f R
un S
3
118
Fig
ure
F.1
6:
Vola
tile
Tem
per
ature
Gra
ph o
f R
un S
4
119
Appendix G. Calculation of Pressure Drop and Actual Heat Transfer
G.1. Pressure Drop
The solution presented here is for Run A4 and „full open‟ gas-exit-valve. The static pressure and gas velocity are 77.608 Pa
(gage) and 0.851 m/s, respectively. Refer to Appendix 4 and 5 for the values of the static pressures and gas velocities, respectively.
The homogeneous, two-phase mode requires that
0.851m sTP G Lv v v
G.1
The densities of each gas component (CO2 and CH4) at the inlet and exit are calculated from Eq. (G.2) and (G.3), respectively.
11
1
p
RT
G.2
22
2
p
RT
G.3
where ρ1 and ρ2 are the inlet and exit densities, respectively. p1 was set equal to p2 because p2 is yet to be solved. After the pressure
drop has been solved, the exit static pressure p2 is determined and inserted back to Eq. (G.3). The new ρ2 is then used to re-compute
the pressure drop in an iterative manner until the solution converges. The values of T1 and T2 were taken from 1:43:00 and are equal to
51.5°C and 30.9°C, respectively; p1 is 77.608 Pa (gage) plus the atmospheric pressure which was 758 mm Hg or 101,058.355 Pa. For
CO2, R is 188.9 J/kg·°C. Hence, the densities are
2
3
1,
77.6081.650 kg m
188.9 51.5 273CO
2
3
2,
77.6081.762 kg m
188.9 30.9 273CO
For CH4, R is 518.2 J/kg·°C. The densities are
4
3
1,
77.6080.601kg m
518.2 51.5 273CH
4
3
2,
77.6080.642 kg m
518.2 30.9 273CH
Run A4 had Pure Green Pellet feedstock which has a composition of 92.08% CO2 and 7.92% CH4. Refer to Table 4.6. The density of
the gas mixture was calculated using Eq. (G.4).
2 2 4 4G CO CO CH CHy y
G.4
where: 2COy = 0.9208
4CHy = 0.0792
At the inlet,
3
1, 0.9208 1.650 0.0792 0.601 1.567 kg mG
and at the exit,
3
2, 0.9208 1.762 0.0792 0.642 1.673 kg mG
120
The average pyrolysis gas density, therefore, is
31.567 1.6731.620 kg m
2G
The absolute viscosities of each gas component are obtained from gas property tables by linear interpolation. Table G.1 shows the
absolute viscosities of CO2 and CH4 at 51.5°C and 30.9°C.
Table G.1: Absolute Viscosities of Pyrolysis Gas Components
Gas Absolute Viscosity, kg/m·s
51.5°C 30.9°C
CO2 1.606 x 10-5 1.513 x 10-5
CH4 1.189 x 10-5 1.131 x 10-5
The absolute viscosity of the gas mixture is calculated using Eq. (G.5).
2 2 4 4G CO CO CH CHy y
G.5
At the inlet,
5 5 5
1, 0.9208 1.606 10 0.0792 1.189 10 1.573 10 kg m sG
and at the exit,
5 5 5
2, 0.9208 1.513 10 0.0792 1.131 10 1.483 10 kg m sG
The average pyrolysis gas viscosity, therefore, is
5 551.573 10 1.483 10
1.528 10 kg m s2
G
The volume of bio-oil collected when the static pressure measurements were taken was 29 ml for the 15-minute sampling interval.
Refer to Figure F.4 in Appendix F. The volume flow rate of bio-oil is estimated using Eq. (G.6).
15 min
boL
VV
G.6
where: Vbo = 15 ml
38 329 ml 1 1L 1m
3.222 10 m s15 min 60 s 1000 ml 1000 L
LV
The flow area occupied by the bio-oil is
Lbo
TP
VA
v
G.7
where: vTP = 0.851 m/s
88 23.222 10
3.786 10 m0.851
boA
The flow area of the pyrolysis gas is
2
4pg i bo i boA A A d A
G.8
where: di = 0.0239 m
121
2 8 4 20.0239 3.786 10 4.486 10 m
4pgA
The volume flow rate of the pyrolysis gas is calculated using Eq. (G.9).
G pg TPV A v
G.9
4 4 34.486 10 0.851 3.818 10 m sGV
The void fraction is calculated using Eq. (G.10).
GG
L G
V
V V
G.10
4
8 4
3.818 100.99992
3.222 10 3.818 10G
The quality is calculated using Eq. (G.11). ρL is the actual density of Pure Green Pellets in Run A4.
1
GG
L
GG G
L
x
G.11
where: ρL = 897.64 kg/m3
ρG = 1.620 kg/m3
εG = 0.99992
1.6200.99992
897.640.955
1.6201 0.99992 0.99992
897.64
x
The density of the two-phase mixture is calculated from Eq. (G.12).
1
G LTP
L Gx x
G.12
where: ρL = 897.64 kg/m3
ρG = 1.620 kg/m3
x = 0.955
31.620 897.641.696 kg m
0.955 897.64 1 0.955 1.620TP
Similarly, the absolute viscosity is calculated from Eq. (G.13). The value of μL is evaluated at 41.2°C for water.
1
G LTP
L Gx x
G.13
where: μL = 6.413 x 10-4 kg/m·s
μG = 1.528 x 10-5 kg/m·s
5 4
5
4 5
1.528 10 6.413 101.598 10 kg m s
0.955 6.413 10 1 0.955 1.528 10TP
122
The mass flux of is calculated from Eq. (G.14), where A is the cross sectional area of the inner tube.
L GTP TP
V Vm
A
G.14
where: LV = 3.222 x 10-8 m3/s
GV = 3.818 x 10-4 m3/s
ρTP = 1.696 kg/m3
A = 4.486 x 10-4 m2
8 4
2
4
3.222 10 3.818 101.696 1.443 kg m s
4.486 10TPm
The Reynolds number is calculated using Eq. (G.15).
Re TP iTP
TP
m d
G.15
where: TPm = 1.443 kg/m2·s
μTP = 1.598 x 10-5 kg/m·s
di = 0.0239 m
5
1.443 0.0239Re 2,158.809
1.598 10TP
For two-phase flow, Re > 2,000 is turbulent.[14] Thus, the friction factor is calculated using Eq. (G.16).
1/40.079ReTP TPf
G.16
1/4
0.079 2,158.809 0.0116TPf
The pressure drop is calculated from Eq. (G.17). L is the distance between the two pressure taps as illustrated in Appendix B; α is the
actual tilt angle of the condenser during the experiment as shown in Appendix B.
2
2sin
TP TP
TP
i TP
f m Lp g L
d
G.17
where: fTP = 0.0116
TPm = 1.443 kg/m2·s
L = 0.738 m
di = 0.0239 m
ρTP = 1.696 kg/m3
α = -25°
g = 9.81 m/s2
2
2 0.0116 1.443 0.7389.81 1.696 0.738 sin 25
0.0239 1.696p
6.067 Pap
The exit pressure is calculated using Eq. (G.18). p2 is inserted back to Eq. (G.3) to recalculate ρ2. Then, the entire solution is
recalculated until the value of Δp converges.
2 1p p p
G.18
where: p1 = 77.608 Pa (gage)
123
2 77.608 6.067 71.541Pa (gage)p
p2 is inserted back to Eq. (G.3) to recalculate ρ2. The first iteration yields a 5.977 x 10-3 % change in the ρ2. This means that the first
solution was already a sufficient estimate of the pressure drop and the mass flux which is used to calculate the actual heat transfer.
For Run A4 and „full open‟ gas-exit-valve, the time when the static pressure and gas velocity were measured is from
1:43:00 to 1:44:50. In between this time duration there are 12 temperature readings at a 10-second interval. Since the gas properties
are highly sensitive to temperature, 12 pressure drop calculations were done corresponding to each of the 12 temperature readings.
The results of these calculations are presented in Table G.2. The same calculations were also conducted for Run A4 „slightly closed‟
and Run A5 both „full open‟ and „slightly closed‟. The summary is shown in Table G.3, G.4, and G.5.
Table G.2: Summary of Pressure Drop for Run A4, „full
open‟
Time, h:min:sec Gas Temperature, °C
Pressure Drop, Pa Inlet Exit
1:43:00 51.5 30.9 6.067
1:43:10 49.6 28.5 6.104
1:43:20 48.6 28.1 6.116
1:43:30 47.8 28 6.124
1:43:40 47.3 27.9 6.129
1:43:50 46.6 27.9 6.134
1:44:00 46.2 27.8 6.139
1:44:10 45.8 27.8 6.142
1:44:20 60.2 51.3 5.551
1:44:30 65.3 55.1 5.495
1:44:40 65.9 40.5 5.871
1:44:50 60.1 30 6.008
Average 5.990
Table G.3: Summary of Pressure Drop for Run A4,
„slightly closed‟
Time, h:min:sec Gas Temperature, °C
Pressure Drop, Pa Inlet Exit
1:45:00 57.6 28.7 5.773
1:45:10 56.3 28.4 5.784
1:45:20 49.2 28.3 5.835
1:45:30 48.3 28.2 5.842
1:45:40 60 31.9 5.731
1:45:50 60.9 32.5 5.720
1:46:00 61.8 33 5.710
1:46:10 63.6 33.7 5.693
1:46:20 64.2 30.9 5.711
1:46:30 64.1 30.1 5.719
1:46:40 60.3 29.3 5.750
1:46:50 57.7 28.3 5.776
Average 5.754
124
G.2. Actual Heat Transfer
In the calculation of the actual heat transferred, the specific heats at constant pressure of the individual gas components
were determined using Eq. (G.19).[11]
2 3a bT cT dTc
M
G.19
where c is the specific heat at constant pressure; T is the gas temperature in Kelvin; and the constants a, b, c, and d are listed in Table
G.6 for CO2 and CH4. The result of Eq. (G.19) is in kJ/kg·K and must be multiplied by 1000 to yield J/kg·K for consistency in units.
Table G.6: Constants for Eq. (G.19)
Gas a b c d M, kg/kmol
CO2 22.26 0.05981 -3.501 x 10-5 7.469 x 10-9 44.010
CH4 19.89 0.05024 1.269 x 10-5 -1.101 x 10-10 16.043
The average temperature which is 41.2°C or 314.2 K is used in the calculation of the specific heat. The specific heats of CO2 and CH4
are respectively
2
2 32 5 9 100022.26 5.981 10 314.2 3.501 10 314.2 7.649 10 314.2
44.010COc
2
859.526 J kg KCOc
4
2 32 5 10 100019.89 5.024 10 314.2 1.269 10 314.2 1.101 10 314.2
16.043CHc
4
2,301.613 J kg KCHc
The specific heat of the gas mixture is calculated using Eq. (G.20)
2 2 4 4G CO CO CH CHc y c y c
G.20
where: 2COy = 0.9208
4CHy = 0.0792
Table G.4: Summary of Pressure Drop for Run A5, „full
open‟
Time, h:min:sec Gas Temperature, °C
Pressure Drop, Pa Inlet Exit
0:15:30 58.8 38.5 5.038
0:15:40 61.8 40.7 5.052
0:15:50 63.6 41.3 5.038
0:16:00 64.6 42 5.028
0:16:10 65.8 42.7 5.016
0:16:20 67.2 43.7 5.002
0:16:30 68.4 44.5 4.991
0:16:40 69.2 45.2 4.982
0:16:50 70 45.6 4.975
0:17:00 70.1 34.9 5.044
Average 5.017
Table G.5: Summary of Pressure Drop for Run A5,
„slightly closed‟
Time,
h:min:sec
Gas Temperature, °C Pressure Drop,
Pa Inlet Exit
0:18:00 67.4 29.4 6.177
0:18:10 67.6 29.9 6.234
0:18:20 67.9 30.7 6.227
0:18:30 68.6 31.4 6.218
0:18:40 69 31.7 6.213
0:18:50 69.4 31.7 6.211
0:19:00 69.6 31.8 6.209
0:19:10 69.8 31.7 6.208
0:19:20 70.2 43.7 6.123
0:19:30 70.3 34.4 6.186
Average 6.201
125
0.9208 859.526 0.0792 2,301.613 973.709 J kg KGc
The two-phase specific heat is calculated from Eq. (G.21). The specific heat of water at 41.2°C is used as the specific heat of bio-oil.
1
G LTP
L G
c cc
xc x c
G.21
where: cG = 973.709 J/kg·K
cL = 4,181.289 J/kg·K
x = 0.955
973.709 4,181.2891,008.266 J kg K
0.955 4,181.289 1 0.955 973.709TPc
The actual heat transfer during 1:43:00 is estimated using Eq. (G.22), where A is the cross sectional area of the inner tube; ΔT is the
difference between the inlet and exit temperature at 1:43:00.
TP TPQ m Ac T
G.22
where: TPm = 1.443 kg/m2·s
A = 4.486 x 10-4 m2
cTP = 1,008.266 J/kg·K
ΔT = 51.5°C - 30.9°C = 20.6°C
41.443 4.486 10 1,008.266 20.6 13.446 WQ
The heat transfer for Runs A4 and A5 during the time when the static pressure readings were taken is summarized in Table G.7, G.8,
G.9, and G.10.
Table G.7: Summary of Heat Transfer for Run A4, „full
open‟
Time, h:min:sec Gas Temperature, °C
Heat Transfer, W Inlet Exit
1:43:00 51.5 30.9 13.446
1:43:10 49.6 28.5 13.827
1:43:20 48.6 28.1 13.450
1:43:30 47.8 28 13.001
1:43:40 47.3 27.9 12.744
1:43:50 46.6 27.9 12.292
1:44:00 46.2 27.8 12.100
1:44:10 45.8 27.8 11.841
1:44:20 60.2 51.3 5.660
1:44:30 65.3 55.1 6.439
1:44:40 65.9 40.5 16.243
1:44:50 60.1 30 19.535
Average 12.548
Table G.8: Summary of Heat Transfer for Run A4,
„slightly closed‟
Time, h:min:sec Gas Temperature, °C
Heat Transfer, W Inlet Exit
1:45:00 57.6 28.7 8.133
1:45:10 56.3 28.4 7.859
1:45:20 49.2 28.3 5.914
1:45:30 48.3 28.2 5.692
1:45:40 60 31.9 7.875
1:45:50 60.9 32.5 7.951
1:46:00 61.8 33 8.055
1:46:10 63.6 33.7 8.349
1:46:20 64.2 30.9 9.317
1:46:30 64.1 30.1 9.520
1:46:40 60.3 29.3 8.705
1:46:50 57.7 28.3 8.276
Average 7.970
126
Table G.9: Summary of Heat Transfer for Run A5, „full
open‟
Time, h:min:sec Gas Temperature, °C
Heat Transfer, W Inlet Exit
0:15:30 58.8 38.5 7.414
0:15:40 61.8 40.7 7.752
0:15:50 63.6 41.3 8.179
0:16:00 64.6 42 8.279
0:16:10 65.8 42.7 8.451
0:16:20 67.2 43.7 8.583
0:16:30 68.4 44.5 8.717
0:16:40 69.2 45.2 8.744
0:16:50 70 45.6 8.883
0:17:00 70.1 34.9 12.929
Average 8.793
Table G.10: Summary of Heat Transfer for Run A5,
„slightly closed‟
Time, h:min:sec Gas Temperature, °C
Heat Transfer, W Inlet Exit
0:18:00 67.4 29.4 5.396
0:18:10 67.6 29.9 5.392
0:18:20 67.9 30.7 5.319
0:18:30 68.6 31.4 5.317
0:18:40 69 31.7 5.330
0:18:50 69.4 31.7 5.387
0:19:00 69.6 31.8 5.401
0:19:10 69.8 31.7 5.444
0:19:20 70.2 43.7 3.772
0:19:30 70.3 34.4 5.124
Average 5.188
127
Appendix H. Recalculation of Double-Pipe Condenser Length
H.1. Bio-oil and Pyrolysis Gas Properties
The calculations presented here are for Seagrass with Binder which has a composition of 90.24% CO2 and 9.76% CH4.
First, the properties of the bio-oil and pyrolysis gas were determined. The properties of the bio-oil were obtained from property tables
of water and listed below according to the condenser zone in which they were applied. In the desuperheating zone the properties were
evaluated at 110°C and 100°C but only the average values are shown in Table H.1, except for the enthalpies. Table H.2 shows the bio-
oil properties applied in the condensing zone and Table H.3 shows the average bio-oil properties applied in the subcooling zone.
Table H.1: Bio-oil Properties Applied in Desuperheating Zone
Property Symbol Numerical Value Unit
Enthalpy at 110°C hi 2,696,200 J/kg
Enthalpy at 100°C hsat 2,676,100 J/kg
Absolute Viscosity μL,des 1.244 x 10-5 kg/m·s
Density ρL,des 0.5816 kg/m3
Thermal Conductivity kL,des 0.02519 W/m·K
Table H.2: Bio-oil Properties Applied in Condensing Zone
Property Symbol Numerical Value Unit
Viscosity of Saturated Steam μv,con 1.227 x 10-5 kg/m·s
Viscosity of Saturated Water μL,con 2.826 x 10-4 kg/m·s
Thermal Conductivity of Saturated Water kL,con 0.682 W/m·K
Density of Saturated Steam ρL,con 0.590 kg/m3
Density of Saturated Water ρv,con 958.320 kg/m3
Latent Heat of Vaporization hfg 2,257,000 J/kg
Specific Heat of saturated Water cL,con 4,211 J/kg·K
Table H.3: Bio-oil Properties Applied in Subcooling Zone
Property Symbol Numerical Value Unit
Absolute Viscosity μL,sub 5.342 x 10-4 kg/m·s
Density ρL,sub 976.709 kg/m3
Thermal Conductivity kL,sub 0.651 W/m·K
Specific Heat at Constant Pressure cL,sub 4,193.165 J/kg·K
The specific heats of the gas components were calculated using Eq. (H.1) and Table H.4. This solution is similar to Eq. (H.19) in
Appendix I.
2 3a bT cT dTc
M
H.1
Table H.4: Constants for Eq. (H.1)
Gas a b c d M, kg/kmol
CO2 22.26 0.05981 -3.501 x 10-5 7.469 x 10-9 44.010
CH4 19.89 0.05024 1.269 x 10-5 -1.101 x 10-10 16.043
At 110°C or 383 K, the specific heat of CO2 and CH4 are respectively,
2
2 32 5 9 100022.26 5.981 10 383 3.501 10 383 7.649 10 383
44.010COc
2
919.138 J kg KCOc
128
4
2 32 5 10 100019.89 5.024 10 383 1.269 10 383 1.101 10 383
16.043CHc
4
2,554.835 J kg KCHc
The specific heat of the gas mixture was calculated using Eq. (H.2). The values of the mass fraction are for Seagrass with Binder
shown in Table 4.6.
2 2 4 4G CO CO CH CHc y c y c
H.2
where: 2COy = 0.9024
4CHy = 0.0976
2COc = 919.138 J/kg·K
4CHc = 2,554.835 J/kg·K
,110 0.9024 919.138 0.0976 2,554.835 1,078.788 J kg KG Cc
The specific heat of the gas mixture at 100°C was also calculated using Eq. (H.1) and (H.2). The results are
2
910.834 J kg KCOc
4
2,517.569 J kg KCHc
,100 1,067.658 J kg KG Cc
For the calculations in the desuperheating zone used the average of cG,110°C and cG,100°C was determined as shown below.
,
1,078.788 1,067.6581,073.223 J kg K
2G desc
The specific heat of the pyrolysis gas at the subcooling zone was calculated by following the same procedures discussed above. The
result of the calculation is shown in Table H.5. The density of the individual gas components were calculated using Eq. (H.3).
p
RT
H.3
where R is the gas constant of the individual gas obtained from gas property tables; p is the actual static pressure measured in the
experiment, as shown in Table 4.3. At 110°C or 383 K and 186.259 Pa (gage) or 101,511.259 Pa (abs), the density of CO2 and CH4 are
respectively,
2
3101,511.2591.403 kg m
188.9 383CO
4
3101,511.2590.511kg m
518.2 383CH
The density of the pyrolysis gas was calculated using Eq. (H.4).
2 2 4 4G CO CO CH CHy y
H.4
where: 2COy = 0.9024
4CHy = 0.0976
2CO = 1.403 kg/m3
4CH = 0.511 kg/m3
129
3
,110 0.9024 1.403 0.0976 0.511 1.316 kg mG C
The density of the gas mixture at 100°C was also calculated using Eq. (H.3) and (H.4). The results are
2
31.441kg mCO
4
30.525 kg mCH
3
,100 1.351kg mG C
For the calculations in the desuperheating zone used the average of ρG,110°C and ρG,100°C was determined as shown below.
3
,
1.316 1.3511.334 kg m
2G des
The density of the pyrolysis gas at the subcooling zone was calculated by following the same procedures discussed above. The result
of the calculation is shown in Table H.5. The absolute viscosity and thermal conductivity of CO2 and CH4 were determined from gas
property tables. The absolute viscosity and thermal conductivity of the gas mixture was calculated using Eq. (H.5) and (H.6),
respectively.
2 2 4 4G CO CO CH CHy y
H.5
2 2 4 4G CO CO CH CHk y k y k
H.6
The values of the properties of the pyrolysis gas determined from the calculations above are summarized in Table H.5 and H.6.
Table H.5: Pyrolysis Gas Properties Applied in Desuperheating Zone
Property Symbol Numerical Value Unit
Specific Heat at Constant Pressure cG,des 1,073.223 kJ/kg
Density ρG,des 1.334 kJ/kg
Absolute Viscosity μG,des 1.791 x 10-5 kg/m·s
Thermal Conductivity kG,des 2.501 x 10-2 W/m·K
Table H.6: Pyrolysis Gas Properties Applied in Subcooling Zone
Property Symbol Numerical Value Unit
Specific Heat at Constant Pressure cG,sub 1,027.952 J/kg·°C
Density ρG,sub 1.505 kg/m3
Absolute Viscosity μG,sub 1.623 x 10-5 kg/m·s
Thermal Conductivity kG,sub 2.160 x 10-2 W/m·K
H.2. Mass Flux
The volume flow rate of bio-oil in the subcooling zone was calculated using Eq. (H.7), where Vbo was taken as 80 ml,
which was the highest recorded bio-oil yield in the 15-minute sampling rate.
,15 min
boL sub
VV
H.7
38 3
,
80 1 1 18.889 10 m s
15 min 60 1000 1000L sub
ml L mV
s ml L
130
The flow area occupied by the bio-oil is
Lbo
TP
VA
v
G.7
where: vTP = 0.125 m/s
87 28.889 10
7.111 10 m0.125
boA
The flow area of the pyrolysis gas is
2
4pg i bo i boA A A d A
G.8
where: di = 0.0239 m
2 7 4 20.0239 7.111 10 4.479 10 m
4pgA
The volume flow rate of pyrolysis gas in the subcooling zone was calculated using Eq. (H.8), where vTP was the recorded velocity
corresponding to the static pressure discussed above as shown in Table 4.4.
,G sub pg TPV A v
H.8
where: Apg = 4.479 x 10-4 m2
vTP = 0.125 m/s
4 5 3
, 4.479 10 0.125 5.599 10 m sG subQ
The void fraction in the subcooling zone was calculated using Eq. (H.9).
,
,
, ,
G sub
G sub
L sub G sub
V
V V
H.9
5
, 8 5
5.599 100.998
8.899 10 5.599 10G sub
The quality in the subcooling zone was calculated using Eq. (H.10). Refer to Table H.3 and H.6 for the values of ρL,sub and ρG,sub.
,
,
,
,
, ,
,
1
G sub
G sub
L sub
sub
G sub
G sub G sub
L sub
x
H.10
1.5050.998
976.7090.492
1.5051 0.998 0.998
976.709
subx
The two-phase density in the subcooling zone was calculated using Eq. (H.11). Refer to Table H.3 and H.6 for the values of ρL,sub and
ρG,sub.
, ,
,
, ,1
G sub L sub
TP sub
sub L sub sub G subx x
H.11
3
,
1.505 976.7093.050 kg m
0.492 976.709 1 0.492 1.505TP sub
131
The mass flux in the subcooling zone was calculated using Eq. (H.12).
, ,
, ,
L sub G sub
TP sub TP sub
V Vm
A
H.12
where: ,L subV = 8.889 x 10-8 m3/s
,G subV = 5.599 x 10-5 m3/s
A = 4.486 x 10-4 m2
8 52
, 4
8.889 10 5.599 103.050 0.381kg m s
4.486 10TP subm
As discussed in Section 3.7.2, the mass flux and quality are constant, that is,
2
, , , 0.381kg m sTP TP des TP con TP subm m m m
0.492sub con desx x x x
H.3. Required Heat Transfer
The following parameters are extensively used in the calculation of the required heat transfer.
2
4 2
0.381 kg m ×s
0.492
4.486 10 m
TPm
x
A
The heat released by the volatiles in the desuperheating zone was calculated using Eq. (H.13).
,1des TP G des desQ m A x h xc T
H.13
where: Δh = hi - hsat = 2,696,200 - 2,676,100 = 20,100 J/kg
ΔTdes = 110 - 100 = 10°C
cG,des = 1,073.223 J/kg·K
40.381 4.486 10 1 0.492 20,100 0.492 1,073.223 10desQ
2.649 WdesQ
The heat released by the bio-oil during condensation was calculated using Eq. (H.14).
1con TP fgQ m x Ah
H.14
where: hfg = 2,257,000 J/kg
40.381 1 0.492 4.486 10 2,257,000conQ
195.949 WconQ
The heat released by the volatiles in the subcooling zone was calculated using Eq. (H.16). First, the specific heat of the two-phase
mixture was solved using Eq. (H.15), where the values of cL,sub and cG,sub are listed in Table H.3 and H.6, respectively.
, ,
,
, ,1
G sub L sub
TP sub
L sub G sub
c cc
xc x c
H.15
,
1,027.952 4,193.1651,666.313 J kg K
0.492 4,193.165 1 0.492 1,027.952TP subc
132
,sub TP TP sub subQ m Ac T
H.16
40.381 4.486 10 1,666.313 100 31subQ
19.668 WsubQ
H.4. Logarithmic Mean Temperature Difference
First, the inlet and exit temperatures of the cooling water at each zone were determined. These temperatures are illustrated
in Figure 3.12; the inlet temperature at the subcooling zone Tw1 was set to 30°C. Tw2, which is the exit temperature at the subcooling
zone, was calculated using Eq. (H.17).
2 1sub
w w
w w
QT T
m c
H.17
where wm is the actual mass flow rate of the cooling water during the experiment which is 0.3 kg/s; cw is the specific heat of water at
30°C which is 4,176 J/kg·K. The value of Qsub was determined in Section H.3.
2
19.66830 30.016 C
0.3 4,176wT
The result of Tw2 was used to calculate Tw3 in Eq. (H.18). The value of Qcon was also determined in Section H.3.
3 2con
w w
w w
QT T
m c
H.18
3
195.94930.016 30.172 C
0.3 4,176wT
Similarly, Tw4 was calculated using Eq. (H.19).
4 3des
w w
w w
QT T
m c
H.19
4
2.64930.172 30.174 C
0.3 4,176wT
The LMTD at the desuperheating, condensing, and subcooling zone were calculated using Eq. (H.20), (H.21), and (H.22),
respectively. Tv,in, Tsat, and Tv,ex are equal to 110°C, 100°C, and 31°C, respectively.
, 4 3
,
, 4
3
ln
v in w sat w
lm des
v in w
sat w
T T T TT
T T
T T
H.20
,
110 30.174 100 30.17274.715 C
110 30.174ln
100 30.172
lm desT
3 2
,
3
2
ln
sat w sat w
lm con
sat w
sat w
T T T TT
T T
T T
H.21
,
100 30.172 100 30.01669.906 C
100 30.172ln
100 30.016
lm conT
133
2 , 1
,
2
, 1
ln
sat w v ex w
lm des
sat w
v ex w
T T T TT
T T
T T
H.22
,
100 30.016 31 3016.238 C
100 30.016ln
31 30
lm subT
H.5. Convection Heat Transfer Coefficients
The convection heat transfer coefficient of the cooling water was calculated as follows. First, the flow area of the annulus
was calculated using Eq. (H.23). The values of Di and do are shown in Appendix B; inner tube diameters were based on the aluminum
tube.
2 2
4i oA D d
H.23
2 2 4 20.0351 0.0254 4.598 10 m
4A
The Reynolds number of the cooling water was calculated using Eq. (H.24). DH was determined from Appendix 1 and μw was
evaluated at 30°C.
Re w Hw
W
m D
A
H.24
where: DH = 0.023 m
μw = 8.030 x 10-4 kg/m·s
ṁw = 0.3 kg/s
4 4
0.3 0.023Re 18,727.60
4.598 10 8.030 10w
For Re > 10,000 and 0.6 < Pr < 100, Eq. (H.25) was used to calculate the Nusselt number of the cooling water. [15] Pr of water was
evaluated at 30°C.
0.8 0.4Nu 0.023Re Pr
H.25
where: Pr = 5.412
Re = 18,727.60
0.8 0.4
Nu 0.023 18,727.60 5.412 118.320
The convection heat transfer coefficient of the cooling water was calculated using Eq. (H.26). kw was evaluated at 30°C.
Nuw ww
H
kh =
D
H.26
where: kw = 0.619 W/m·K
2118.320 0.619= = 3,179.614 W m K
0.023wh
134
The convection heat transfer coefficient at the desuperheating zone was calculated as follows. The absolute viscosity of the
two-phase mixture was calculated using Eq. (H.27).
, ,
,
, ,1
G des L des
TP des
L des G desx x
H.27
where: x = 0.492
μL,des = 1.244 x 10-5 kg/m·s
μG,des = 1.791 x 10-5 kg/m·s
5 5
5
, 5 5
1.791 10 1.244 101.464 10 kg m s
0.492 1.244 10 1 0.492 1.791 10TP des
The Reynolds number at the desuperheating zone was calculated using Eq. (H.28).
,
,
Re TP iTP des
TP des
m d
H.28
where: TPm = 0.381 kg/m2·s
di = 0.0239 m
μTP,des = 1.464 x 10-5 kg/m·s
, 5
0.381 0.0239Re 622.426
1.464 10TP des
The thermal conductivity of the two-phase mixture was calculated using Eq. (H.29).
, ,
,
, ,1
G des L des
TP des
L des G des
k kk
xk x k
H.29
where: x = 0.492
kL,des = 2.519 x 10-2 W/m·K
kG,des = 2.501 x 10-2 W/m·K
2 2
2
, 2 2
2.501 10 2.519 102.510 10 W m K
0.492 2.519 10 1 0.492 2.501 10TP desk
Since the Reynolds number indicates that the flow is laminar, the Nusselt number of the flow is equal to 3.66 for constant wall
temperature, that is,
Nu 3.66v,des i
des
TP,des
h d= =
k
H.30
Rearranging Eq. (H.30) yields
-2
23.66 2.501×10
= = 3.844 W m K0.0239
v,desh
The convection heat transfer coefficient at the condensing zone was calculated as follows. The Reynolds number of the
flow was calculated using Eq. (H.31).
,
Re TP icon
v con
m d
H.31
where: TPm = 0.381 kg/m2·s
di = 0.0239 m
μv,con = 1.227 x 10-5 kg/m·s
135
5
0.381 0.0239Re 742.777
1.227 10con
Since Recon < 35,000, Eq. (H.32) was used to calculate the convection heat transfer coefficient during condensation.
1/4 1/43sin 0.68
0.555L,con L,con v,con L,con fg L,con sat i
v,con
L,con i sat i
ρ ρ ρ g α k h + c T Th =
μ d T T
H.32
Eq. (H.32), however, contains a variable which is unknown, which is Ti. As derived from Appendix A, Eq. (H.33) was used to
calculate Ti.
1/4 1/43
2
sin ln0.68 ln0.555
22
ln
t io w w,con
L,con L,con vcon L,con v,in i i o ifg L,con sat i o i
itL,con i sat i t
o w
o i
k Td h T +
ρ ρ ρ g α k T T d d dh + c T T d dT = +
kμ d T -T kd h +
d d
H.33
The numerical values of the variables in Eq. (H.33) are shown in Table H.7. Tube diameters are shown in Appendix B.
Table H.7: Values of Variables in Eq. (H.33)
Variable Numerical Value Unit Variable Numerical Value Unit
ρL,con 958.320 kg/m3 hfg 2,257,000 J/kg
ρv,con 0.590 kg/m3 hw 3,179.614 W/m2·K
μL,con 2.826 x 10-4 kg/m·s Tw,con 30.164 °C
α 20 deg Tsat 100 °C
kL,con 0.6816 W/m·K Tv,in 110 °C
kt 204 W/m·K g 9.81 m/s2
cL,con 4,211 J/kg·K
Applying bisection method to Eq. (H.33) yields,
= 74.964 °CiT
Substituting Ti back to Eq. (H.32) yields the convection heat transfer coefficient at the condensing zone,
2
, W m Kv conh =5,974.601
The convection heat transfer coefficient at the subcooling zone was calculated as follows. The absolute viscosity of the
two-phase mixture was calculated using Eq. (H.34).
, ,
,
, ,1
G sub L sub
TP sub
L sub G subx x
H.34
where: x = 0.492
μL,sub = 5.342 x 10-4 kg/m·s
μG,sub = 1.623 x 10-5 kg/m·s
5 4
5
, 4 5
1.623 10 5.342 103.196 10 kg m s
0.492 5.342 10 1 0.492 1.623 10TP sub
The Reynolds number at the subcooling zone was calculated using Eq. (H.35).
,
,
Re TP iTP sub
TP sub
m d
H.35
where: TPm = 0.381 kg/m2·s
136
di = 0.0239 m
μTP,sub = 3.196 x 10-5 kg/m·s
, 5
0.381 0.0239Re 285.149
3.196 10TP des
The thermal conductivity of the two-phase mixture was calculated using Eq. (H.36).
, ,
,
, ,1
G sub L sub
TP sub
L sub G sub
k kk
xk x k
H.36
where: x = 0.492
kL,sub = 0.651 W/m·K
kG,sub = 2.160 x 10-2 W/m·K
2
2
, 2
2.160 10 0.6514.241 10 W m K
0.492 0.651 1 0.492 2.160 10TP subk
Since the Reynolds number indicates that the flow is laminar, the Nusselt number of the flow is equal to 3.66 for constant wall
temperature, that is,
Nu 3.66v,sub i
sub
TP,sub
h d= =
k
H.37
Rearranging Eq. (H.37) yields
-2
23.66 4.241×10
= = 6.495 W m K0.0239
v,subh
H.6. Condenser Length
The computed values of Q, ΔTlm, and hv for each zone were substituted to Eq. (H.38) to solve the length for each zone.
ln1 1
Δ 2
o i
lm i v t o w
d dQL = + +
π T d h k d h
H.38
For the desuperheating zone,
ln1 1
Δ 2
o idesdes
lm,des i v,des t o w
d dQL = + +
π T d h k d h
where: Qdes = 2.649 W
ΔTlm,des = 74.715°C
hv,des = 3.844 W/m2·K
ln 0.0254 0.02392.649 1 1= + +
74.715 0.0239 3.844 2 204 0.0254 3,179.614desL
π
= 0.123 mdesL
For the condensing zone,
ln1 1
Δ 2
o iconcon
lm,con i v,con t o w
d dQL = + +
π T d h k d h
where: Qcon = 195.949 W
ΔTlm,con = 69.906°C
137
hv,con = 5,974.601 W/m2·K
ln 0.0254 0.0239195.949 1 1= + +
69.906 0.0239 5,974.601 2 204 0.0254 3,179.614conL
π
= 0.017 mconL
For the subcooling zone,
1 1
Δ 2
o isubsub
lm,sub i v,sub t o w
ln d dQL = + +
π T d h k d h
where: Qcon = 19.668 W
ΔTlm,con = 16.238°C
hv,con = 6.495 W/m2·K
ln 0.0254 0.023919.668 1 1= + +
π 16.238 0.0239 6.495 2 204 0.0254 3,179.614subL
= 2.488 msubL
The total condenser length is the sum of the lengths of each zone.
des con subL= L +L +L
H.39
= 0.123+0.017 + 2.488 = 2.629 mL
H.7. Pressure Drop
Eq. (H.40) was used to calculate the pressure drop in each zone.
2
2sin
TP TP
TP
i TP
f m Lp g L
d
H.40
where α is the ideal tilt angle of the condenser which is -20° from the horizontal. Since the Reynolds number at all three zones was
found to be laminar, Eq. (H.41) was used to calculate the friction factors in each zone.
16
ReTP
TP
f
H.41
The density of the two-phase mixture was calculated using Eq. (H.42).
1
G LTP
L Gx x
H.42
For the desuperheating zone,
,
,
16
ReTP des
TP des
f
where: ReTP,des = 622.426
2
,
162.571 10
622.426TP desf
The two-phase density is,
, ,
,
, ,1
G des L des
TP des
L des G desx x
138
where the values of ρG,des and ρL,des are shown in Table H.1 and H.5, respectively.
3
,
1.334 0.58160.805 kg m
0.492 0.5816 1 0.492 1.334TP des
The pressure drop at the desuperheating zone is
2
,
,
,
2sin
TP des TP des
des TP des des
i TP des
f m Lp g L
d
222 2.571 10 0.381 0.1239.81 0.805 0.123 sin 20
0.0239 0.805desp
0.380 Padesp
The pressure at the desuperheating zone exit is
2 1 186.259 0.380 185.879 Pa (gage)desp p p
This is equal to 101,510.879 Pa (abs). The absolute value was substituted back to Eq. (H.3) to recalculate the exit density of the
individual pyrolysis gas components. The result was a 1.897 x 10-4 % change in the value of ρG,des, which is a negligible change.
Therefore, recalculation of the gas properties and its dependent parameters was not necessary.
For the condensing zone,
,
,
16
ReTP con
TP con
f
where: ReTP,con = 742.777
2
,
162.154 10
742.777TP conf
From the assumption stated in Section 3.7.5 about the condensing zone, the bio-oil enters the condensing zone in vapor-phase and
leaves in liquid-phase. Two-phase density, therefore, is
, ,
,2
v con L con
TP con
where the values of ρv,con and ρL,con are shown in Table A10.7.
3
,
0.590 958.320479.455 kg m
2TP con
The pressure drop at the condensing zone is,
2
,
,
,
2sin
TP con TP con
con TP con con
i TP con
f m Lp g L
d
22 2
22 5.611 10 0.381 1.743 10
9.81 479.455 1.743 10 sin 200.0239 479.455
conp
28.038 Paconp
The pressure at the condensing zone exit is
3 2 185.879 28.038 157.841Pa (gage)conp p p
139
For the subcooling zone,
,
,
16
ReTP sub
TP sub
f
where: ReTP,sub = 285.149
2
,
165.611 10
284.149TP subf
The two-phase density is,
, ,
,
, ,1
G sub L sub
TP sub
L sub G subx x
where the values of ρG,sub and ρL,sub are shown in Table H.3 and H.6, respectively.
3
,
1.505 976.7093.050 kg m
0.492 976.709 1 0.492 1.505TP sub
The pressure drop at the subcooling zone is,
2
,
,
,
2sin
TP sub TP sub
sub TP sub sub
i TP sub
f m Lp g L
d
222 5.611 10 0.381 2.4889.81 3.050 2.488 sin 20
0.0239 3.050subp
26.027 Pasubp
The pressure at the subcooling zone exit is
4 3 157.841 26.027 131.814 Pa (gage)subp p p
This is equal to 101,456.814 Pa (abs). The absolute value was substituted back to Eq. (H.3) to recalculate the exit density of the
individual pyrolysis gas components at the subcooling zone. The result was a 0.03 % change in the value of ρG,des, which is also
negligible. Therefore, recalculation of the gas properties and its dependent parameters was not necessary.
The calculations presented above were also performed for the other five feedstock listed in Table 4.6 in Section 4.5. Table
H.8 and H.9 shows a summary of the required condenser length and pressure drop, respectively, of each feedstock in Table 4.6.
Table H.8: Summary of Required Condenser Length
Marine Florae Feedstock Condenser Length, cm
Desuperheating Zone Condensing Zone Subcooling Zone Total
Seagrass w/ Binder 12.3 1.7 248.8 262.8
Pure Red Pellets 12.0 1.7 239.6 253.3
Red Raw 12.0 1.7 237.9 251.6
Green Raw 12.2 1.7 246.8 260.7
Pure Green Pellets 12.3 1.7 249.6 263.6
Brown w/ Binder 12.4 1.7 249.9 264.0
140
Table H.9: Summary of Pressure Drop
Marine Florae Feedstock Pressure Drop, Pa (gage)
Desuperheating Zone Condensing Zone Subcooling Zone Total
Seagrass w/ Binder 0.38 28.04 26.03 54.45
Pure Red Pellets 0.34 28.04 23.67 52.06
Red Raw 0.34 28.04 23.37 51.75
Green Raw 0.36 28.04 25.20 53.60
Pure Green Pellets 0.37 28.04 25.98 54.39
Brown w/ Binder 0.37 28.04 26.09 54.50
141
Definition of Terms
Annular Space – the hollow space between two concentric tubes of different diameter;
the space in the double-pipe condenser where the cooling water flows
Biomass – plant and animal material, especially agricultural waste products, used as a
source of fuel
Condenser – a type of heat exchanger specifically designed to
Cooling Water – the water that flows outside the inner-tube, in the annular space, of the
condenser for the purpose of cooling the volatiles
Desuperheating – the process of reducing the temperature of superheated steam (bio-oil)
down to 100°C (at atmospheric pressure)
Feed Port – the part of the reactor where the marine florae feedstock enter the reactor
Feedstock – raw material required for a certain process; the marine florae required for
the pyrolysis process
Gas Mixture – the same as pyrolysis gas
Gas-Exit-Pipe – the pipe in the reactor that was designed as the channel where the
volatiles exit the reactor; the designed interface of the reactor and condenser
Heat Exchanger – an apparatus where two or more fluids exchange heat for a specific
purpose
Homogeneous Mixture – a fluid system composed of more than component that is well
mixed so that the properties of the mixture are uniform in the entire system
Hopper – integral part of the reactor feed port
Layer A of the Reactor – the part of the reactor where the last thermocouple probe was
placed before the volatiles exit the reactor.
LMTD – Logarithmic Mean Temperature Difference
Quality – the ratio of the mass of gas/vapor in the system to the total mass of the system;
the mass of pyrolysis gas that flows in the condenser
Pipe/Tube – in general, flow sections of circular cross section are referred to as pipes.
Small diameter pipes are usually referred to as tubes; in this text pipe and tube are
used interchangeably
142
Pyrolysis Reactor – apparatus where the pyrolysis reaction takes place; where the
marine florae feedstock are heated in the absence of oxygen
Residence Time – time duration that the liberated volatiles remain in the reactor; time
required to complete the pyrolysis reaction
Run – a single experiment trial involving only one type of feedstock and either of the two
condensers
Subcooling – the process of reducing the temperature of liquid bio-oil from 100°C to a
lower temperature without freezing it
Two-Phase – the flow of fluids where there are more than one phase present; the bio-oil
and pyrolysis gas flowing in the condenser at the same time
Void Fraction – the space occupied by the pyrolysis gas in the two-phase flow inside the
condenser
Volatiles – more commonly known as volatile matter; refers to the mixture of the bio-oil
and pyrolysis gas
Zone – a portion of the condenser given a specific heat exchange duty: desupereheating,
condensing, and subcooling
143
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