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.9: Bio-oil Volume Collected for Run S5 ......................................................101

Table E.10: Bio-oil Volume Collected for Run S6 ....................................................101



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.9: Volatile Temperature Graph of Run S5 ..................................................111

Figure F.10: Volatile Temperature Graph of Run S6 ................................................112







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


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





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


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


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.


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


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


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


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


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


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)



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)



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)



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


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


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


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


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


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


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


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


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


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


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


%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


%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


%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).


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).


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


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‟



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)


G.20


4CHy = 0.0792

Table G.4: Summary of Pressure Drop for Run A5, „full

open‟



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,


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


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‟


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,




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‟



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,




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


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


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.


H.2


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).


H.4


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.


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


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


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


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


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


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


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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development of a condenser for marine florae pyrolysis reactor

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