Development of a Condenser for Marine Florae Pyrolysis Reactor

Development of a Condenser for the DevelopedMarine 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 CO 2 and CH 4 . The bio-oil was also observed to stick to the walls of the condenser. Among the two condensers, the aluminum condenser was easier to clean and had less oil that stick to its walls. The required condenser length was recalculated by incorporating the determined rate of bio- oil yield and pyrolysis gas components to the calculations. Also, the two-phase flow of the volatiles was considered in the recalculation. The recalculated length was found to be impractically too long for a double-pipe condenser. Other types of condensers, e.g. shell- and-tube, are suggested for future studies. Engr. Edwin A. Carcasona, PhD, PME Thesis Adviser iii TABLE OF CONTENTS Acknowledgement .............................................................................................................. i Abstract .............................................................................................................................. ii List of Tables ................................................................................................................... vii List of Figures ................................................................................................................... ix Chapter 1. Problem Setting and Background .................................................................1 1.1. Introduction .............................................................................................................1 1.2. Statement of the Problem ........................................................................................2 1.3. Significance of the Study ........................................................................................3 1.4. Objectives ...............................................................................................................4 1.5. Scope and Limitation ..............................................................................................4 1.5.1. Condenser Design ..........................................................................................5 1.5.2. Sources and Types of Marine Florae Feedstock ............................................6 1.6. Theoretical Background ..........................................................................................6 1.6.1. Pyrolysis of Marine Florae ............................................................................6 1.6.2. Condensation Phenomenon............................................................................8 1.6.3. Flow Rates .....................................................................................................8 1.6.4. Conservation Laws ........................................................................................9 1.6.5. Heat Transfer ...............................................................................................10 1.6.6. Gas Mixtures ................................................................................................13 1.6.7. Dimensionless Numbers ..............................................................................14 1.6.8. Homogeneous Two-Phase Model ................................................................15 Chapter 2. Review of Related Literature.......................................................................18 2.1. Pyrolysis Of Biomass............................................................................................18 2.1.1. Bio-oil ..........................................................................................................19 2.1.2. Liquid Collection .........................................................................................20 2.1.3. Pyrolysis Gas ...............................................................................................20 2.2. Double-Pipe versus Other Types of Condensers ..................................................21 2.2.1. Shell-and-Tube.............................................................................................21 iv 2.2.2. Spiral-Tube ..................................................................................................22 2.2.3. Plate-Fin .......................................................................................................22 2.2.4. Gasketed Plate..............................................................................................23 2.2.5. Spiral Plate ...................................................................................................24 2.2.6. Direct Contact ..............................................................................................24 2.3. Condensers Used in Pyrolysis...............................................................................25 2.3.1. Unapumnuk (1999) .....................................................................................25 2.3.2. Mudolodu (2002) ........................................................................................25 2.3.3. Jih (1982) ....................................................................................................26 2.3.4. Añora (2010) ...............................................................................................26 2.4. Condensation of Mixtures .....................................................................................27 2.5. Research Gap ........................................................................................................28 Chapter 3. Methodology ..................................................................................................29 3.1. Introduction ...........................................................................................................29 3.2. Condenser Design Process ....................................................................................30 3.2.1. Required Heat Transfer ................................................................................31 3.2.2. Convection Heat Transfer Coefficient .........................................................34 3.2.3. Logarithmic Mean Temperature Difference ................................................36 3.2.4. Heat Transfer Area.......................................................................................36 3.3. Marine Florae Collection and Preparation ............................................................37 3.4. Installation of Centrifugal Blower ........................................................................37 3.5. Experiment Set-up and Procedure ........................................................................39 3.5.1. Equipment Preparation ................................................................................41 3.5.2. Cooling Water Flow Calibration..................................................................42 3.5.3. Fluid Temperature Measurement .................................................................43 3.5.4. Periodic Oil Collection and Measurement ...................................................44 3.5.5. Static Pressure Measurement .......................................................................45 3.5.6. Gas Velocity Measurement ..........................................................................46 3.5.7. Gas Collection for Gas Chromatography.....................................................48 3.6. Condenser Evaluation ...........................................................................................48 3.6.1. Cleanability ..................................................................................................48 v 3.6.2. Pressure Drop ...............................................................................................49 3.6.3. Actual Heat Transferred...............................................................................53 3.7. Recalculation of the Double-Pipe Condenser Length...........................................54 3.7.1. Properties of Bio-oil and Pyrolysis Gas .......................................................55 3.7.2. Mass Flux .....................................................................................................57 3.7.3. Required Heat Transfer ................................................................................58 3.7.4. Logarithmic Mean Temperature Difference ................................................59 3.7.5. Convection Heat Transfer Coefficients .......................................................60 3.7.6. Length of the Condenser ..............................................................................60 3.7.7. Pressure Drop ...............................................................................................61 Chapter 4. Results and Discussion .................................................................................62 4.1. Designed and Fabricated Double-Pipe Condenser ...............................................62 4.2. Temperature of Condenser Fluids.........................................................................63 4.2.1. Temperature of Volatiles .............................................................................63 4.2.2. Temperature of Cooling water .....................................................................66 4.3. Static Pressure and Gas Velocity ..........................................................................69 4.4. Bio-oil Yield .........................................................................................................71 4.4.1. Effect of Blower on Bio-oil Yield ...............................................................72 4.4.2. Bio-oil Leakage............................................................................................74 4.4.3. Black Viscous Liquid...................................................................................74 4.5. Pyrolysis Gas ........................................................................................................76 4.5.1. Components .................................................................................................76 4.5.2. Estimate of Pyrolysis Gas Yield ..................................................................77 4.6. Condenser Performance ........................................................................................77 4.6.1. Condenser Material ......................................................................................78 4.6.2. Pressure Drop ...............................................................................................79 4.6.3. Actual Heat Transferred...............................................................................79 4.7. Results of Recalculation of The Condenser Length .............................................80 4.7.1. Comparison of Initial Calculation and Recalculation ..................................80 4.7.2. Effect of Flow Velocity ...............................................................................81 4.7.3. Effect of Thermal Conductivity of Condenser Tube ...................................81 vi 4.7.4. Effect of Cooling Water ...............................................................................82 Chapter 5. Conclusions and Recommendations............................................................84 5.1. Conclusions ...........................................................................................................84 5.2. Recommendations .................................................................................................85 Appendices ........................................................................................................................88 Appendix A. Calculation of Initial Condenser Design ................................................88 A.1. Required Heat Transfer ..................................................................................88 A.2. Logarithmic Mean Temperature Difference ..................................................89 A.3. Convection Heat Transfer Coefficients..........................................................90 A.4. Condenser Length ..........................................................................................92 Appendix B. Fabricated Condenser Parts and Assembly ............................................93 B.1. Dimensions and Parts .....................................................................................93 B.2. Condenser Accessories ...................................................................................94 B.3. Thermocouple Probes and Pressure Taps.......................................................95 B.4. Condenser Tilt Angle .....................................................................................96 Appendix C. Cooling Water Flow Rate Measurements ..............................................97 Appendix D. Static Pressure and Gas Velocity Measurements ...................................98 D.1. Static Pressure Measurements ........................................................................98 D.2. Gas Velocity Measurements ..........................................................................99 Appendix E. Pyrolysis Products ................................................................................100 E.1. Periodic Bio-oil Volume Measurement ........................................................100 E.2. Bio-oil and Pyrolysis Gas Yield ...................................................................102 E.3. Product Composition and Residence Time from Añora (2010) ..................103 Appendix F. Volatile Temperature Graph .................................................................104 F.1. Volatile Temperature Graph with Plotted Periodic Bio-oil Yield ................104 F.2. Volatile Temperature Graph without Plotted Periodic Bio-oil Yield ...........115 Appendix G. Calculation of Pressure Drop and Actual Heat Transfer ......................119 G.1. Pressure Drop ...............................................................................................119 G.2. Actual Heat Transfer ....................................................................................124 Appendix H. Recalculation of Double-Pipe Condenser Length ................................127 vii H.1. Bio-oil and Pyrolysis Gas Properties ...........................................................127 H.2. Mass Flux .....................................................................................................129 H.3. Required Heat Transfer ................................................................................131 H.4. Logarithmic Mean Temperature Difference ................................................132 H.5. Convection Heat Transfer Coefficients........................................................133 H.6. Condenser Length ........................................................................................136 H.7. Pressure Drop ...............................................................................................137 Definition of Terms ........................................................................................................141 Bibliography ...................................................................................................................143 LIST OF TABLES Table 1.1: Types of Marine Florae Feedstock ...............................................................6 Table 3.1: Experiment Runs ........................................................................................41 Table 3.2: Necessary Bio-oil Properties ......................................................................56 Table 3.3: Necessary Gas Properties ...........................................................................56 Table 4.1: Cooling Water Temperature Reading for Run A2 ......................................67 Table 4.2: Cooling Water Inlet and Volatile Exit Temperatures for Run S1 ..............68 Table 4.3: Inlet Static Pressure ....................................................................................69 Table 4.4: Gas Velocity in the Condenser Inner-Tube ................................................70 Table 4.5: Collected Bio-oil for Run A7 .....................................................................72 Table 4.6: Component Percentage of Pyrolysis Gas....................................................76 Table A.1: Values of Variable in Eq. (23) ..................................................................92 Table B.1: List of Parts ................................................................................................93 Table B.2: Inner-Tube Actual Dimensions ..................................................................94 Table B.3: Outer-Tube Actual Dimensions .................................................................94 Table C.1: Mass Flow Rate for Fully Open .................................................................97 Table C.2: Mass Flow Rate for One Valve-Turn .........................................................97 Table C.3: Mass Flow Rate for Two Valve-Turns ......................................................97 viii Table C.4: Mass Flow Rate for Three Valve-Turns ....................................................97 Table C.5: Mass Flow Rate for Four Valve-Turns ......................................................97 Table D.1: Manometer Reading for Run A4, „full open‟ ............................................98 Table D.2: Manometer Reading for Run A4, „slightly close‟ .....................................98 Table D.3: Manometer Reading for Run A5, „full open‟ ............................................98 Table D.4: Manometer Reading for Run A5, „slightly close‟ .....................................98 Table D.5: Gas Exit Velocity .......................................................................................99 Table D.6: Velocity Inside Inner-Tube ........................................................................99 Table E.1: Bio-oil Volume Collected for Run A1 .....................................................100 Table E.2: Bio-oil Volume Collected for Run A2 .....................................................100 Table E.3: Bio-oil Volume Collected for Run A3 .....................................................100 Table E.4: Bio-oil Volume Collected for Run A4 .....................................................100 Table E.5: Bio-oil Volume Collected for Run A5 .....................................................100 Table E.6: Bio-oil Volume Collected for Run A6 .....................................................100 Table E.7: Bio-oil Volume Collected for Run A7 .....................................................101 Table E.8: Bio-oil Volume Collected for Run A8 .....................................................101 Table E.9: Bio-oil Volume Collected for Run S5 ......................................................101 Table E.10: Bio-oil Volume Collected for Run S6 ....................................................101 Table E.11: Bio-oil Volume Collected for Run S7 ....................................................101 Table E.12: Bio-oil Volume Collected for Run S8 ....................................................102 Table E.13: Mass of Marine Florae Feedstock and Pyrolysis Products ....................102 Table E.14: Mass Percentage of Pyrolysis Products ..................................................102 Table E.15: Density of Bio-oil ...................................................................................103 Table E.16: Mass Percentage and Residence Time for Green Algae ........................103 Table E.17: Mass Percentage and Residence Time for Red Algae............................103 Table E.18: Mass Percentage and Residence Time for Brown Algae .......................103 Table E.19: Mass Percentage and Residence Time for Seagrass...............................103 Table G.1: Absolute Viscosities of Pyrolysis Gas Components ................................120 Table G.2: Summary of Pressure Drop for Run A4, „full open‟ ................................123 Table G.3: Summary of Pressure Drop for Run A4, „slightly close‟ .........................123 Table G.4: Summary of Pressure Drop for Run A5, „full open‟ ................................124 ix Table G.5: Summary of Pressure Drop for Run A5, „slightly close‟ .........................124 Table G.6: Constants for Eq. (G.19) .........................................................................124 Table G.7: Summary of Heat Transfer for Run A4, „full open‟ ................................125 Table G.8: Summary of Heat Transfer for Run A4, „slightly close‟ .........................125 Table G.9: Summary of Heat Transfer for Run A5, „full open‟ ................................126 Table G.10: Summary of Heat Transfer for Run A5, „slightly close‟ .......................126 Table H.1: Bio-oil Properties Applied in Desuperheating Zone................................127 Table H.2: Bio-oil Properties Applied in Condensing Zone......................................127 Table H.3: Bio-oil Properties Applied in Subcooling Zone ......................................127 Table H.4: Constants for Eq. (H.1) ...........................................................................127 Table H.5: Pyrolysis Gas Properties Applied in Desuperheating Zone .....................129 Table H.6: Pyrolysis Gas Properties Applied in Subcoolnig Zone ............................129 Table H.7: Values of Variables in Eq. (H.33) ...........................................................135 Table H.8: Summary of Required Condenser Length ...............................................139 Table H.9: Summary of Pressure Drop ......................................................................140 LIST OF FIGURES Figure 1.1: Double-Pipe Heat Exchanger ......................................................................2 Figure 1.2: Average Mass Loss Curve with respect to Time (without binder) .............7 Figure 1.3: Average Mass Loss Curve with respect to Time (with binder)...................7 Figure 1.4: Diagram of Double-Pipe Counter-Flow Heat Exchanger .........................11 Figure 1.5: Temperature Profile for Counter-Flow Heat Exchanger ...........................12 Figure 2.1: Shell-and-Tube Condenser ........................................................................21 Figure 2.2: Plate-Fin Condenser ..................................................................................22 Figure 2.3: Gasketed Plate Heat Exchanger ................................................................23 Figure 2.4: Experiment Set-up of Mudulodu (2002) ..................................................25 Figure 2.5: Experiment Set-up of Jih (1982) ..............................................................26 Figure 2.6: Experiment Set-up of Añora (2010) .........................................................27 Figure 3.1: Study Flow ................................................................................................29 x Figure 3.2: Condenser Design Flow Chart ..................................................................31 Figure 3.3: Gas Escaping through the Feed Port of the Reactor..................................38 Figure 3.4: RPM and Air Velocity Measurement ........................................................38 Figure 3.5: Retrofitted Centrifugal Blower..................................................................39 Figure 3.6: Schematic of Experiment Set-up ...............................................................39 Figure 3.7: Actual Experiment Set-up without Manometer ........................................40 Figure 3.8: Insulated Condenser ..................................................................................42 Figure 3.9: Installed Condenser ...................................................................................42 Figure 3.10: Thermocouple Datalogger .......................................................................43 Figure 3.11: Condenser with Thermocouple Probes ...................................................44 Figure 3.12: Bio-oil Collection and Storage ................................................................44 Figure 3.13: Static Pressure Measurement Set-up .......................................................45 Figure 3.14: Inclination Positioning Instruments ........................................................45 Figure 3.15: Gas-Exit-Valve Positions ........................................................................47 Figure 3.16: Uro-bag filled with Pyrolysis Gas ...........................................................48 Figure 3.17: Temperature Profile.................................................................................55 Figure 3.18: Conservation of Mass in the Condenser ..................................................58 Figure 4.1: Condenser Length .....................................................................................62 Figure 4.2: Condenser Tilt Angle ................................................................................62 Figure 4.3: Volatile Temperature Graph of Run A1 ....................................................63 Figure 4.4: Temperature Rise while Blower was Turned Off for Run A4 ..................64 Figure 4.5: Volatile Exit and Cooling Water Inlet Temperatures for Run S1 .............65 Figure 4.6: Cooling water Exit Temperature for Run S1 ............................................67 Figure 4.7: Fan-System Curve .....................................................................................70 Figure 4.8: Collected Bio-oil .......................................................................................71 Figure 4.9: Volatile Temperature Graph......................................................................73 Figure 4.10: Bio-oil Leakage Plotted in Volatile Temperature Graph ........................74 Figure 4.11: Unrecovered Black Viscous Liquid ........................................................74 Figure 4.12: Collected Black Viscous Liquid ..............................................................75 Figure 4.13: Black Viscous Liquid Residue ................................................................75 Figure 4.14: Flame from Pyrolysis Gas .......................................................................76 xi Figure 4.15: Comparison of Stainless and Aluminum Condensers .............................78 Figure 4.16: Aluminum Condenser .............................................................................78 Figure 4.17: Flow Velocity, Condenser Length, Pressure Drop ..................................81 Figure 4.18: Cooling Water Convection Coefficient and Condenser Length..............82 Figure B.1: Aluminum Condenser ...............................................................................93 Figure B.2: Stainless Condenser ..................................................................................93 Figure B.3: Exploded View of the Condenser .............................................................94 Figure B.4: Adapter .....................................................................................................94 Figure B.5: Static Pressure Tap ...................................................................................95 Figure B.6: Position of Thermocouple Probes .............................................................95 Figure B.7: Position of Static Pressure Taps ...............................................................95 Figure B.8: Condenser Tilt Angle ................................................................................96 Figure F.1: Volatile Temperature Graph of Run A1 ..................................................104 Figure F.2: Volatile Temperature Graph of Run A2 ..................................................105 Figure F.3: Volatile Temperature Graph of Run A3 ..................................................106 Figure F.4: Volatile Temperature Graph of Run A4 ..................................................106 Figure F.5: Volatile Temperature Graph of Run A5 ..................................................107 Figure F.6: Volatile Temperature Graph of Run A6 ..................................................108 Figure F.7: Volatile Temperature Graph of Run A7 ..................................................109 Figure F.8: Volatile Temperature Graph of Run A8 ..................................................110 Figure F.9: Volatile Temperature Graph of Run S5 ..................................................111 Figure F.10: Volatile Temperature Graph of Run S6 ................................................112 Figure F.11: Volatile Temperature Graph of Run S7 ................................................113 Figure F.12: Volatile Temperature Graph of Run S8 ................................................114 Figure F.13: Volatile Temperature Graph of Run S1 ................................................115 Figure F.14: Volatile Temperature Graph of Run S2 ................................................116 Figure F.15: Volatile Temperature Graph of Run S3 ................................................117 Figure F.16: Volatile Temperature Graph of Run S4 ................................................118 1 CHAPTER 1 PROBLEM SETTING AND BACKGROUND 1.1. Introduction Energy is a resource that the modern society cannot live without. However, studies have shown that the conventional method of harnessing energy, e.g. burning of fossil fuels, is literally steadily killing the environment. To prevent further destruction of the environment attention has been given to renewable energy sources. One such renewable energy resource, that is the topic of this study, is biomass, particularly marine florae or more commonly known as seaweeds. Earlier studies regarding the use of marine florae as an energy resource have been conducted by Baring, et al (2009) [3] and Añora (2010) [1] and were able to present promising results. Añora studied the extraction of useful fuel products from marine florae by means of a method known as pyrolysis. This was replicated by Esgana (2011) [13] on a much larger scale. Pyrolysis is the heating of the biomass in the absence of oxygen to produce solid, liquid, and gaseous end products such as carbonaceous char, bio-oil, and pyrolysis gases, respectively. [25][26] During pyrolysis reaction volatiles are released from the biomass. These volatiles are composed of the condensable and noncondensable components, which are the bio-oil and pyrolysis gas, respectively. Esgana‟s study was to develop a marine florae pyrolysis reactor that was capable of pyrolyzing much larger quantities of marine florae than in Añora‟s study. In any pyrolysis system the pyrolysis reactor is coupled with a condenser for separating the condensable from the noncondensable component of the volatiles, which is discussed in references [1], [16], [20], [24]. The purpose of the present study was to develop the condenser for the marine florae pyrolysis reactor. The pyrolysis experiment of the present study was done simultaneously with Esgana‟s experiment. Since pyrolysis product yield is also dependent on the heat rate and reactor conditions, the results of this study is applicable only for the reactor developed by Esgana or other reactors of the similar specifications. At present, there are no recognized design methods and most work has been empirical and specific to the characteristics of the feedstock being processed. 2 Commercial liquids recovery processes are usually proprietary and may be specific to individual feedstock, reactor configurations and products. [7] Being the first condenser designed for the developed marine florae pyrolysis reactor, the simplest type of condenser was seen as the best starting point. The type of condenser chosen was the double-pipe heat exchanger, shown in Figure 1.1, because it is easy to fabricate, maintain, and its tubular construction allows it to be easily “scaled up” to shell-and-tube if greater heat exchange duties are required. Also, a tubular condenser is effective in separating the oil from the pyrolysis gas. [9] Cleanability of the condenser was a major concern because the volatiles may contain solid char particles due to carry-over from the pyrolysis reactor. [25] The carry-over char was also observed in the present study. It was also observed from the study of Añora (2010) that the bio-oil deposited to the walls of the reactor which then required frequent cleaning. The performance of the initial design was analyzed: flaws and problems were identified and improvements were suggested. Figure 1.1: Double-Pipe Heat Exchanger 1.2. Statement of the Problem Design of condensers requires that the properties of both the cooling fluid and the vapor to be condensed be known. However, at the beginning of this study the researcher did not have data regarding the properties of the marine florae volatiles. Thus, assumptions of the composition and properties of marine florae volatiles were required to be able to come up with the initial condenser design. The assumptions were based on literatures on the pyrolysis of other types of biomass (mostly wood). [8][25][26] Properties of the volatiles of other biomass could be far different from that of marine florae which 3 could result in the failure of the condenser operation: the calculated heat transfer area might be insufficient, thus the required amount of heat rejection might not be attained and the condensable component might not be condensed. Another problem that existed in the design of the condenser was the presence of the noncondensable component. Condensation of mixtures with noncondensable gases is a complicated phenomenon. [14] Some assumptions were necessary in order to simplify the calculations but, in the process, sacrificed the accuracy of the solution. Another factor that was considered in condenser design was the material. In the case of corrosive fluids, one might need to use expensive corrosion-resistant materials such as stainless steel or even titanium. [10] In this study two types of materials were used to construct the condenser, and were visually inspected for any signs of corrosion. The two types of material were also tested for cleanability since it was observed in the experiment of Añora (2010) that the bio-oil sticks to the walls of the distilling flask, which was used as the pyrolysis reactor, and the glass condenser. This could cause fouling in the condenser which would decrease the effectiveness. After the experiment, the researcher inspected which condenser material had less bio-oil that stick to it and which was more easily cleaned. 1.3. Significance of the Study Literatures support that the volatiles obtained from biomass are good candidates as alternative fuels. [2][21][23] Since the volatiles are composed of the condensable and noncondensable components, a condenser is needed to condense the condensable component and separate it from the noncondensable component. There are many parameters that must be considered in condenser design. Some of these parameters are properties of the marine florae volatiles, volatile flow rate, and condenser operating conditions. Most of these parameters were unknown prior to the present study. This study designed and fabricated a condenser based only on assumptions of the design parameters. The purpose of the fabricated condenser was to expose an actual condenser to the operating conditions encountered in the pyrolysis of marine florae using the reactor developed by Esgana. In this manner some of the design parameters were 4 revealed during the experiment and after analysis of the acquired data. The present study also investigated if the condenser material had any significant effect. There is concrete theory on condenser design; however, good knowledge of the environment in which the condenser will operate must be known to be able to achieve the optimum design. Also, operating conditions vary with different systems in which the condenser is used; hence, each condenser design is unique to its own system. In pyrolysis, the amount of volatiles and volatile flow differ with type of feedstock and reactor design. Since the reactor used in the present study was a new design developed by Esgana [13] , the actual operating conditions and the behavior of the volatiles is yet to be known. Hence, investigation regarding the actual operating conditions was necessary. Problems that arose in the experiment were identified and solutions were suggested. The results of this study can be used as a bench mark for future condenser designs for the developed marine florae pyrolysis reactor. [13] 1.4. Objectives - To design and fabricate two double-pipe condensers, one made from aluminum and the other from stainless steel, for the marine florae pyrolysis reactor. - To evaluate the performances of the fabricated condensers. - To determine the percent amount of bio-oil and pyrolysis gas that can be extracted and collected from a given marine florae feedstock. - To reconstruct the condenser design methodology based on the collected data from the experiment. 1.5. Scope and Limitations The scope of this study was to make a simple design, that is, the two-phase flow of the volatiles was not considered, and to fabricate the condenser for the marine florae pyrolysis reactor developed by Esgana. The condenser was used to separate the condensable from the noncondensable component of the marine florae volatiles. The researcher investigated how the condenser performed during the experiment. The amount of volatiles extracted from the feedstock was also monitored, especially the bio-oil to aid in the analysis of the condenser performance. Specifications/parameters for future 5 condenser design were derived based on the performance of the fabricated condenser and other data obtained from the experiment. This study was not concerned with the operation of the pyrolysis reactor. However, data on the reactor temperatures were referred to whenever appropriate. Other parts or components directly related to the operation of the reactor were not included in this study. 1.5.1. Condenser Design The condenser was fabricated using materials that were affordable and available in the local market. Two double-pipe condensers were fabricated; one with an aluminum inner tube and the other with stainless steel inner tube. Very little was known about the marine florae volatiles, especially its thermophysical properties which were necessary parameters in designing the condenser. The properties of the volatiles were assumed based on literatures on other biomass pyrolysis. The assumptions are discussed in Section 3.2 and in Appendix A together with the details of the calculation. The flow configuration that was tested in the experiment was counter flow only. The volatiles flowed inside the inner tube and the cooling water in the annular space between the inner and outer tubes. Counter flow was chosen over parallel flow because of its superior heat transfer capability. [10] The new design methodology was revised based on the obtained data from the experiment. Only the condenser length was recalculated. The recalculation is discussed in 3.7 and Appendix H. This study was focused only on the thermal design of the condenser and only little detail was given to the bio-oil collection system. 6 1.5.2. Sources and Types of Marine Florae Feedstock The types of marine florae that were used as feedstock for the pyrolysis reactor are Red, Brown, Green algae, and Seagrass only. Only drifted marine florae were used in the experiment and collection was limited to the province of Cebu only. The types of marine florae feedstock that were used in this study are listed in Table 1.1. Table 1.1: Types of Marine Florae Feedstock Green algae Brown algae Red algae Seagrass Pelletized without binder Pelletized with binder Pelletized without binder Pelletized with binder Non-pelletized Pelletized without binder Non-pelletized Pelletized without binder 1.6. Theoretical Background The equations discussed in this Section are presented in their most basic form. The equations take other forms through the discussions depending on the situation in which they were used: type of fluid, flow condition (laminar of turbulent), geometry of the flow area, etc. The applications of the equations presented here are discussed in Chapters 3 and 4. 1.6.1. Pyrolysis of Marine Florae The study of Añora (2010) [1] had proven the possibility of extracting condensable liquid (bio-oil) and combustible gas (pyrolysis gas) from marine florae. There were three events that happened during pyrolysis experiments. First, from 0 minute to 4 minutes, there was no change in the weight of the loaded sample pellets. Next event was the rapid change of sample pellets‟ weight at 4 minutes to 35 minutes. A rapid bio-oil production was observed from 4 minutes to 25 minutes. Also, start of pyrolysis gas production was observed from 7 minutes to 15 minutes. Finally, a slow decrease of sample pellets‟ weight was observed from 35 minutes to 1 hour. It was suspected that pyrolysis reaction has stopped since no production of bio-oil nor pyrolysis gas was observed. 7 Figure 1.2: Average Mass Loss Curve with respect to Time (without binder) [1] Figure 1.3: Average Mass Loss Curve with respect to Time (with binder) [1] Figure 1.2 and 1.3 shows the average mass loss curves with respect to time for pellets with and without binder, respectively. The residence time of the experiment, from start to the end of the experiment, are summarized in Appendix E. The percent product compositions of bio-oil and pyrolysis gas are also shown in Appendix E. These data were used to estimate for the mass flow rate of the volatiles which was then used to solve for the required heat transfer in the initial design of the condenser. 8 1.6.2. Condensation Phenomenon When condensing at low velocity, tube condensation is better in down-flowing inclined tubes than either vertical or horizontal tubes. This is because the layer of condensate in the bottom is quite thick in a horizontal tube, so that a small inclination in the direction of flow results in more rapid condensate flow and much thinner condensate layer. Vertical tubes are not usually as good as inclined ones because the condensate layer is uniform around the tube; better heat transfer is obtained when the condensate layer is nonuniformly distributed. The optimal inclination for condensation is about 20°. [14] 1.6.3. Flow Rates Basic forms of the mass and volume flow rate [11] are equations shown in Eq. (1.1) and (1.2), respectively. These equations were mainly used to solve the flow rates of the volatiles and the cooling water. Other forms of Eq. (1.1) and (1.2) were also used and discussed in Chapters 3 and 4. A = = A m m vA t µ ( ) 1.1 = = m V vA µ ( ) 1.2 where: ṁ = mass flow rate, kg/s V = volume flow rate, m 3 /s ρ = density of fluid, kg/m 3 v = velocity of flow, m/s A = flow area, m 2 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 out m m m A = ÷ ( ) 1.3 where: Δm = net change in mass within a system, kg m in = total mass entering a system, kg m out = total mass leaving a system, kg Eq. (1.3) can also be expressed in the rate form as [11] = ÷ in out dm m m dt ( ) 1.4 where: dm/dt = rate of change of mass within a system, kg/s m in = total mass flow rate into a system, kg/s m out = 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 out m 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 absorbed Q Q ( ) 1.6 where: Q released = heat released by the hot fluid, W Q absorbed = 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] = A = A p Q m h mc T ( ) 1.7 where: Q = heat transfer, W ṁ = mass flow rate, kg/s Δh = change in enthalpy, J/kg c p = 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] A lm T Q R = ( ) 1.8 where: ΔT lm = 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). ( ) ln 1 1 2 o i i i o o d d R= + + Ah πkL A h ( ) 1.9 where: A i = inner surface area of the inner tube, m 2 A o = outer surface area of the inner tube, m 2 d i = inside diameter of the inner tube, m d o = outside diameter of the inner tube, m L = length of the heat exchanger, m k = thermal conductivity of heat exchanger material, W/m· K h i = convection coefficient of fluid inside the inner tube, W/m 2 · K h o = convection coefficient of fluid in the annular space, W/m 2 · K Figure 1.4: Diagram of Double-Pipe Counter-Flow Heat Exchanger 12 The LMTD is ( ) 1 2 1 2 Δ Δ ln Δ Δ lm T T T T T A ÷ = ( ) 1.10 where ΔT 1 and ΔT 2 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 k h L = ( ) 1.11 where: Nu = Nusselt number, dimensionless k = thermal conductivity of the fluid, W/m· K L c = characteristic length, m The Nusselt number is discussed in Section 1.6.7. The characteristic length L c is equal to the tube diameter for flow inside tubes; in annular flow the hydraulic diameter is used in place of L c . 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/ 4 3 0.555 ( = ( ( ¸ ¸ v fg sat i ρ ρ - ρ g'k h' h μd T -T ( ) 1.12 where: ( ) 0.68 fg fg w sat i h' h c T -T = + sin g' = g α ρ = density of liquid film, kg/m 3 ρ v = density of vapor, kg/m 3 μ = absolute viscosity of liquid film, kg/m· s k = thermal conductivity of liquid film, W/m· K h fg = 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 T sat = saturation temperature of vapor, °C T i = inner surface temperature of condenser wall, °C g = acceleration due to gravity, 9.81 m/s 2 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: X mixture = represents any property of the mixture X j = property of a single gas component j y j = 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 c ρvL μ = ( ) 1.14 where: Re = Reynolds number, dimensionless ρ = density of the fluid, kg/m 3 v = velocity of the flow, m/s L c = 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/m 2 · K k = thermal conductivity of the fluid, W/m· K L c = 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 TP v v v ( ) 1.16 where: v G = velocity of the gas-phase, m/s v L = velocity of the liquid-phase, m/s v TP = 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 c µ µ c c µ µ ( ) 1.17 where the void fraction ε G is = + G G L G V V V c ( ) 1.18 where: ρ G = density of the gas-phase, kg/m 3 ρ L = density of the liquid -phase, kg/m 3 G V = volume flow rate of the gas-phase, m 3 /s L V = volume flow rate of the liquid-phase, m 3 /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 L TP L G x x µ µ µ µ µ = + ÷ ( ) 1.19 ( ) 1 = + ÷ G L TP L G x x µ µ µ µ µ ( ) 1.20 where: μ G = absolute viscosity of the gas-phase, kg/m· s μ L = absolute viscosity of the liquid-phase, kg/m· s The mass flux of the mixture is given by Eq. (1.21). - | | + = | \ . G L TP TP V V m A µ ( ) 1.21 where: - TP m = mass flux, kg/m 2 · s A = cross sectional area of the flow, m 2 From the given relations above the pressure drop can be predicted from Eq. (1.22). Eq. (1.22) neglects the accelerational pressure gradient. ( ) 2 2 sin TP TP TP TP f m L p g L D µ o µ - A = ÷ ( ) 1.22 where: f TP = 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 Re TP TP f = ( ) 1.23 for laminar flow (Re < 2,000), or 1/4 0.079Re ÷ = TP TP f ( ) 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 ṁ bo and ṁ pg 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 m m t ( ) 3.1 ( ) % pg pg m m 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,ex Q m h h h c T T ( = ÷ + + ÷ ¸ ¸ ( ) 3.3 where the enthalpies h i and h sat were obtained from steam tables; h fg and c w are the latent heat of vaporization and specific heat of water; T sat is the saturation temperature of water 33 at 1 atm; T v,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 ΔT 2 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,ex Q ym c T T = ÷ ¿ ( ) 3.4 where y was the assumed mass fraction of each gas component; c p is the specific heat of each gas component; T v,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 pg Q 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 0.68 0.555 3 w w v fg w sat i w v w i sat i ρ ρ ρ h c T T g sinα k h μ d T -T ( ( ÷ + ÷ ¸ ¸ = ( ( ¸ ¸ ( ) 3.6 where ρ w is the density of water; ρ v is the density of steam at 105°C; μ w is absolute viscosity of water; k w is the thermal conductivity of water; c w is the specific heat of water; h fg is the latent heat of vaporization of water; T sat is the saturation temperature of water; T i is the inner surface temperature of condenser wall; d i 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 pg i v v v i v m m d m Aμ πd μ 4 + = = ( ) 3.7 where A is the cross sectional area of the inner tube; μ v 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 z 1 was the estimated height of the upper reservoir. 2 1 2 v 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 v w was computed from v 2 and is discussed in Appendix A; D H is the hydraulic diameter of the annulus; the cooling water properties, i.e. ρ w and μ w , 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 0.4 Nu 0.023 Re Pr w w w = ( ) 3.10 Nu w w w H k h D = ( ) 3.11 where Pr w and k w 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 Q T T ρ Vc = + ( ) 3.12 where T w,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 2 ln ln v,in w,ex v,ex w,in lm v,in w,ex v,ex w,in T T T T T T T T T T T T T ÷ ÷ ÷ A ÷A A = = A A | | ÷ | ÷ \ . ( ) 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. ( ) ln 1 1 Δ 2 o i lm i v t o w d d Q L 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 (2 nd ) 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 µ = ( ) 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 ex A v v A = ( ) 3.16 m Av µ = ( ) 3.17 where v is the gas velocity inside the inner tube; ṁ is the mass flow rate; A ex is the cross sectional area of the gas-exit-valve of the condenser; v ex 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 L v v v = = ( ) 3.18 where v TP is the two-phase flow velocity that was used in the calculation of the pressure drop; v G was the estimated pyrolysis gas velocity; v L 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% CO 2 and 7.92% CH 4 ; run A5 used a Red Raw feedstock that has a composition of 75% CO 2 and 23% CH 4 . 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 CO 2 and CH 4 , were calculated from the ideal gas equation shown in Eq. (3.19). p RT µ = ( ) 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 ρ 1 and ρ 2 represent the gas densities at the inlet and exit of the condenser, respectively. For ρ 1 , p 1 is the inlet static pressure that was measured in the experiment. Since there was no data on the exit static pressure p 2 , p 1 was used to solve ρ 2 . After the pressure drop was solved, the exit static pressure p 2 was determined and inserted back to the original solution of ρ 2 . The new ρ 2 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 CO 2 and CH 4 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 CO 2 and CH 4 , thus, the properties of the gas mixture were calculated from Eq. (3.20) and (3.21). 2 2 4 4 G CO CO CH CH y y µ µ µ = + ( ) 3.20 2 2 4 4 G CO CO CH CH y y µ µ µ = + ( ) 3.21 where ρ G and μ G are the density and absolute viscosity of the gas mixture (pyrolysis gas), respectively; 2 CO y and 4 CH y are the mass fractions of each gas component; 2 CO µ and 4 CH µ are the densities of each gas component computed from Eq. (3.19); 2 CO µ and 4 CH µ 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 bo L V V = ( ) 3.22 G TP V Av = ( ) 3.23 where L V and G V are the volume flow rates of the bio-oil and pyrolysis gas, respectively; V bo 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; v TP 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. G G L G V V V c = + ( ) 3.24 1 G G L G G G L x µ c µ µ c c µ | | | \ . = | | ÷ + | \ . ( ) 3.25 where ε G and x are the void fraction and quality, respectively; L V and G V are the volume flow rates of the bio-oil and pyrolysis gas, respectively; ρ L and ρ G 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 L TP L G x x µ µ µ µ µ = + ÷ ( ) 3.26 ( ) 1 G L TP L G x x µ µ µ µ µ = + ÷ ( ) 3.27 where ρ TP and μ TP are the two-phase density and absolute viscosity, respectively; x is the quality; ρ L and ρ G are the densities of the bio-oil and pyrolysis gas, respectively; μ L and μ G are the absolute viscosities of the bio-oil and pyrolysis gas, respectively. The mass flux TP m - was then determined from using Eq. (3.28). L G TP TP V V m A µ - | | + = | \ . ( ) 3.28 Reynolds number Re TP was computed from Eq. (3.29) to determine if the flow was laminar or turbulent. Re TP i TP TP m d µ - = ( ) 3.29 where TP m - is the mass flux; d i is the inner diameter of the inner tube; μ TP 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 Re TP TP f = ( ) 3.30 1/4 0.079Re TP TP f ÷ = ( ) 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 2 sin TP TP TP i TP f m L p g L d µ o µ - A = ÷ ( ) 3.32 2 1 p p p = ÷A ( ) 3.33 where Δp is the pressure drop; f TP 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; ρ TP and TP m - are the density and mass flux of the two-phase mixture, respectively; p 2 is the exit pressure; p 1 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 4 G CO CO CH CH c y c y c = + ( ) 3.34 54 where c G is the specific heat at constant pressure of the pyrolysis gas; 2 CO c and 4 CH c are the specific heats of each gas component; 2 CO y and 4 CH y 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 L TP L G c c c xc x c = + ÷ ( ) 3.35 where c TP is the specific heat of the two-phase mixture; c G and c L 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 TP Q m Ac T - = A ( ) 3.36 where Q is the heat transferred; c TP 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; TP m - 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 CO 2 and CH 4 . 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 CO 2 and CH 4 that were needed in the solution are listed in Table 3.3. Table 3.3: Necessary Gas Properties Condition Properties at 110°C specific heat at constant pressure, absolute viscosity, density, thermal conductivity at 100°C specific heat at constant pressure, absolute viscosity, density, thermal conductivity at 31°C specific heat at constant pressure, absolute viscosity, density, thermal conductivity The specific heat, absolute viscosity, and thermal conductivity were obtained from gas property tables as a function of temperature only. Their variation with pressure was not 57 considered because of the small pressure gradient of the flow which is discussed later. The density was calculated using Eq. (3.20) in Section 3.6.2. The pyrolysis gas was treated as a homogeneous gas mixture and its properties as a whole were calculated using equations similar to Eq. (3.21) and (3.22) in Section 3.6.2. 3.7.2. Mass Flux Eq. (3.22) and (3.23) in Section 3.6.2 were used to calculate the volume flow rate of the bio-oil and pyrolysis gas, respectively; the void fraction, quality, and mass flux were calculated using Eq. (3.24), (3.25), and (3.28), respectively. The volume flow rate of the bio-oil was estimated based on the 15-minute sampling rate of the collected bio-oil volume discussed in Section 3.5.4. In the calculation of the volume flow rate of bio-oil in Eq. (3.22), the volume of the bio-oil was set to the highest recorded volume of bio-oil in a 15-minute duration which was 80 ml; see Figure F.8 in Appendix F. This is equivalent to a volume flow rate of 5.33 ml/min. It is evident from the experiment results shown in Appendix F that the volume of bio-oil collected vary in every 15-minute interval sampling rate. If the condenser is designed to condense the highest amount of bio-oil then it is certain to condense the lesser amounts. The value of the velocity that was used to calculate the volume flow rate of the pyrolysis gas was the actual velocity measured in the experiment. The flow parameters discussed above were all calculated based on the actual experiment condition, wherein the condenser seemed to be just a subcooling heat exchanger because of undesired condensation of the bio-oil prior to the condenser inlet. This is discussed further in Section 4.6. The volume flow rates and void fraction may be different in the desuperheating and condensing zones but the mass flux and quality must be constant to satisfy the principle of conservation of mass. Disregarding the small amount of oil that sticks to the condenser walls, the mass leaving the condenser must be equal to the mass entering the condenser. That is, the mass of the volatiles that leave the subcooling zone is equal to the mass that enter the subcooling zone, which is the same mass that leave the condensing zone, and so on. 58 The statement above is illustrated in Figure 3.18 where m represents the mass of the volatiles. Figure 3.18: Conservation of Mass in the Condenser However, in reality the mass entering the condenser is greater than the mass leaving the condenser because of the small amounts of bio-oil that adhere to the walls of the condenser and accumulate over time. More about this oil is discussed in Section 4.4. The consequence of the assumption that the mass flux is constant is that the calculation of the total required heat transfer will be overestimated. But since the researcher did not have the means to determine the change in mass flux, the calculations were done with constant mass flux. 3.7.3. Required Heat Transfer In the desuperheating zone, the equation used in the calculation of the required heat transfer was slightly different from that used in Section 3.6.3. In the case of superheated steam the specific enthalpy must be used instead of the specific heat in the calculation of heat transfer as shown in Eq. (3.37). The heat released by the pyrolysis gas was calculated using Eq. (3.38). ( ) 1 L TP Q m x A h - = ÷ A ( ) 3.37 G TP G Q m xAc T - = A ( ) 3.38 where Q L is the amount of heat released by bio-oil (assumed as steam) during desuperheating from 110°C to 100°C; Q G is the heat released by the gas from 110°C to 100°C; TP m - 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 G Q 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 h fg is the latent heat of vaporization of water at 1 atm. ( ) 1 con TP fg Q 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 TP Q m Ac T - = A ( ) 3.41 where c TP 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 T w1 , T w2 , T w3 , and T w4 . T w1 and T w2 are the inlet and exit temperatures in the subcooling zone, respectively; T w2 is also the inlet temperature in the condensing zone and T w3 is the exit temperature; T w3 is also the inlet temperature in the desuperheating zone and T w4 is the exit temperature. In the calculations T w1 was set to 30°C. T w2 , T w3 , and T w4 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. k TP , μ TP , and ρ TP , 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, ΔT lm , and h v in each zone. The total length of the condenser was calculated using Eq. (3.42), where L des , L con , and L sub are the lengths of the desuperheating, condensing, and subcooling zones, respectively. des con sub L L L L = + + ( ) 3.42 3.7.7. Pressure Drop Since it was observed from the test run, as discussed in Section 3.4, that the blower is essential equipment in the pyrolysis set-up, the pressure drop inside the condenser must be estimated so that the appropriate size of the blower can be selected. The same procedure and equations presented in Section 3.6.2 for calculating the pressure drop were used. Different flow conditions were solved and the results are compared in Section 4.7. The details of all the calculation presented here in Section 3.7 are presented in Appendix H. 62 CHAPTER 4 RESULTS AND DISCUSSION 4.1. Designed and Fabricated Double-Pipe Condenser The designed condenser lengths, indicated by the symbol L in Figure 4.1, were 78.1 cm and 99.9 cm for aluminum and stainless condenser, respectively. However, due to fabrication errors the actual fabricated lengths were 52 cm and 61 cm for aluminum and stainless condenser, respectively. The full detail on the designed condenser is shown in Appendix B. TC1, TC2, TC3, and TC4 in Figure 4.1 indicate the slots where the thermocouple probes were inserted. TC1 and TC2 measured the inlet and exit temperatures, respectively, of the volatiles. TC3 and TC4 measured the inlet and exit temperatures, respectively, of the cooling water. Refer to Appendix B for the exact positions of the thermocouple probes. Figure 4.1: Condenser Length The designed tilt angle of the condenser was also not realized because of the orientation of the blower with respect to the reactor. The actual tilt angle during the experiment was 25° with respect to the horizontal as shown in Figure 4.2. The procedure for measuring the angle is presented in Appendix B. Figure 4.2: Condenser Tilt Angle 63 4.2. Temperature of Condenser Fluids 4.2.1. Temperature of Volatiles The inlet temperature of the volatiles in the condenser was much lower than the temperature recorded in Layer A of the reactor [13] , shown in Figure 4.3. This meant that there was heat rejection that caused a temperature drop between the reactor and condenser. The temperature drop was due to the relatively long distance that the volatiles had to travel before getting to the condenser inlet. Because the gas-exit-pipe of the reactor leading to the condenser was initially at a lower temperature than the volatiles in the reactor, the volatiles gave off heat as they pass through the pipe. The gas-exit pipe was insulated, so, theoretically, there had to be a certain time when the pipe temperature will attain thermal equilibrium with the volatiles. However, still before the condenser inlet was the blower which was not insulated. Continuous heat rejection to the environment occurred in the blower which led to the large temperature drop and condensation of bio-oil in the blower that was observed because of the leakage. Bio-oil leakage is discussed in Section 4.4.2. Figure 4.3: Volatile Temperature Graph of Run A1 The volatile inlet temperature in the condenser was also constantly changing with time, also shown in Figure 4.3. This meant that the mass flow rate of the volatiles changed with time. When the mass flow rate was high the temperature drop was less 64 because of the increased heat capacity of the volatiles. It was observed in the experiment that when the blower was turned on the volatile inlet temperature was higher than when the blower was turned off, also shown in Figure 4.3. Turning the blower on increased the velocity of the volatiles, thus, increased the mass flow rate. The bio-oil yield was also observed to be high when the blower was turned on, discussed in Section 4.4.1. There was also an instance when the volatile inlet temperature was relatively high even when the blower was turned off. Unlike the sudden increase in temperature when the blower was turned on, the temperature rise was gradual as shown in Figure 4.4. Figure 4.4: Temperature Rise while Blower was Turned Off for Run A4 High volumes of bio-oil were collected during this phenomenon in the runs that it occurred, which meant that there was an increase in the mass flow of the volatiles. Since the blower was turned off, the only reason for the increased mass flow is that the conditions inside reactor changed, increasing the rate of devolatilization. However, this is beyond the scope of the present study. Based on the observations above, the volatile inlet temperature could be considered as an indirect indication of the bio-oil yield, however, only if the reactor temperatures remain fairly constant since bio-oil yield also depends on the reactor temperature. [25] High inlet temperatures had high yield and low inlet temperatures had relatively lower yield. The volatile temperature graph with indicated bio-oil yield is presented in Appendix F. 65 The time of some of the instances when the blower was turned on was recorded and then located on the temperature graph of the corresponding run. Refer to Figure 4.3. During run A1, it was recorded that the blower was turned on from 1:34-1:35. During time 1:34-1:35 there was an abrupt rise of the inlet and exit temperatures of the volatiles, and then a sudden drop after time 1:35 when the blower was turned off. There were numerous accounts of this event throughout the entire experiment. It was concluded that the abrupt rise in temperature was an indication that the blower was turned on. This relationship between the temperature and blower was useful in the analysis of the bio-oil yield which is discussed in Section 4.4.1. When the blower was not turned on the exit temperature of the volatiles was almost equal to the inlet temperature of the cooling water, considering the tolerance of the thermocouple reading discussed in Section 4.2.2. This was observed for all the runs. However, when the blower was turned on, there was an abrupt increase in the exit temperature of the volatiles as shown in Figure 4.5. However, when the gas-exit-valve was in the „slightly close‟ position while the blower was turned on, the rise in exit temperature was not very high. There were even instances when the value of the exit temperature was near the cooling water temperature when the blower was turned on and the gas-exit-valve was in the „slightly close‟ position. Figure 4.5: Volatile Exit and Cooling Water Inlet Temperatures for Run S1 66 The increase in exit temperature was mainly due to the increase in mass flow rate and can easily be explained mathematically, as shown below. From heat balance, for an ideal system with no heat loss, the heat released by the volatiles equals the heat absorbed by the cooling water. Heat Released by Volatiles = Heat Absorbed by Cooling Water ( ) , , ÷ = A v v v in v ex w w w m c T T m c T ( ) 4.1 where ṁ v and c v are the mass flow rate and specific heat of the volatiles, respectively; T v,in and T v,ex are inlet and exit temperatures of the volatiles, respectively; ṁ w and c w are the mass flow rate and specific heat of the cooling water; ΔT w is the change in temperature of the cooling water. Rearranging Eq. (4.1) yields the solution for the exit temperature of the volatiles. , , A = ÷ w w w v ex v in v v m c T T T m c ( ) 4.1 When the blower was turned on, ṁ v and T v,in are increased while c v , ṁ w , c w , and ΔT w remain fairly constant. The increase in ṁ v when the blower was turned on was large enough to significantly affect the 2 nd term in the right side of Eq. (4.1) above, thus increasing T v,ex . However, when the gas-exit-valve was in the „slightly close‟ position, the increase in ṁ v was not large enough to affect T v,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 T v,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 v C C ( ) 4.2 which is equal to , - >> w w TP TP sub m c m Ac 68 ( )( ) ( )( )( ) 4 0.3 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/m 2 · s (v TP = 0.851 m/s), the heat capacity rates are 622.224 J s K 0.754 J s K w v C 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 OP a , which corresponds to a static pressure p a and velocity v a . The point where the fan curve and the system curve (b) – gas valve „slightly close‟ – intersect is the operating point OP b , which corresponds to a static pressure p b and velocity v b . Figure 4.7 confirms the values of the static pressure and gas velocity measurements discussed above. That is, the static pressure at OP a is lower than OP b , and the gas velocity at OP a is higher than OP b . 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 of pyrolysis gas feedstock char bio oil = ÷ ÷ - ( ) 4.3 This estimate of the pyrolysis gas also includes the mass of black viscous liquid that was not measured, as discussed in Section 4.4.3, and the residue left in the reactor. The average mass percentage of pyrolysis gas based on the feedstock was 20.13%. This is higher than the pyrolysis gas yield from the study of Añora (2010) which was 13.09%. Again, the reason might be the configuration of the pyrolysis reactor, as discussed in Section 4.4.3. The list of the pyrolysis gas yield of the different feedstock in every run is tabulated in Table E.2 in Appendix E. 4.6. Condenser Performance During the entire experiment the condenser seemed to be just a subcooling heat exchanger because of undesired condensation of the bio-oil prior to the condenser inlet. The oil leak observed in the blower was evidence that there was condensation there. Certainly, the water content [8] of the bio-oil was condensed in the blower. Since the bio- oil was considered to have a condensing temperature equal to water because the condensing temperatures of other bio-oil components were not known, only the subcooling section of the fabricated condenser was analyzed. The desuperheating and condensing sections were not analyzed. In spite of this, the undesired condensation in the blower provided some clue to the condensation temperature of the bio-oil which was discussed in Section 4.4.2. The effectiveness of the condenser was not calculated because of the discrepancy between the cooling water inlet temperature and volatile exit temperature, as discussed in Section 4.2.2. Also, the actual overall heat transfer coefficient was not calculated because of the same reason. 78 4.6.1. Condenser Material As discussed in Section 4.4.3, there was a black viscous liquid that was observed to adhere to the condenser walls that could cause fouling. When the condensers were disassembled from the pyrolysis set-up to be cleaned after the experiment, less black viscous liquid was observed in the walls of the aluminum condenser than in the stainless condenser as shown in Figure 4.15. The task of cleaning the condensers was mainly to remove the black viscous liquid from the condenser walls. The aluminum condenser was easier to clean than the stainless condenser. Figure 4.15: Comparison of Stainless and Aluminum Condensers The inner tube of the aluminum condenser was almost entirely free of the black viscous liquid after spraying water through the inner tube. Figure 4.16 shows the walls of the inner tube of the aluminum condenser before and after spraying with tap water. The stainless condenser, on the other hand, needed a test tube brush to clean the inner tube. Figure 4.16: Aluminum Condenser Also, there were no obvious indications of corrosion in both materials after a total of 40.27 hours and 37.32 hours of operation for the aluminum and stainless condenser, b) Aluminum Condenser a) Stainless Condenser b) After spraying with water a) Before spraying with water 79 respectively. However, it was later discovered, through further literature survey, that bio- oils can be corrosive to aluminum due to the presence of acetic acid and formic acid. [21] Aside from the physical observations discussed above, aluminum tube is also cheaper than stainless. Aluminum costs P70/m while stainless costs P140/m in the local market. In terms of thermal requirements, aluminum seems to have the advantage over stainless steel because of its higher thermal conductivity, 204 W/m· K for aluminum and 16.3 W/m· K for stainless. [15] However, based on the recalculation discussed in Section 4.7.3, the thermal conductivity of the condenser has little effect on the condenser length. There was black viscous liquid that also stick to the GI pipe fittings, which was even harder to clean. Galvanized iron is not recommended for condenser material. 4.6.2. Pressure Drop The calculations for the pressure drop are shown in Appendix I and the values of the pressure drop at different gas velocities are tabulated in Table I.2 to I.5. The pressure drop in the condenser was very small, an average of 5.990 Pa for Run A4 and „full open‟ gas-exit-valve. The maximum pressure drop determined was from Run A5 and „slightly closed‟ gas-exit-valve position, which was 6.201 Pa. The small pressure drop was not enough to significantly affect the value of the exit density of the pyrolysis gas as discussed in Appendix I. The change in exit density of the pyrolysis gas was 5.977 x 10 -3 %. This means that the initial estimate of the two-phase density, mass flux, and heat transfer, which is discussed in Section 4.6.3, were sufficiently accurate. 4.6.3. Actual Heat Transferred The actual heat transferred, calculated in Appendix I.2, is much lower than the calculated value in the initial design from Eq. (A.5) of Appendix A. In the initial design calculation, the parameters were manipulated to yield the maximum possible heat transfer, as discussed in Section 3.2.1. The required heat transfer calculated in Appendix A was 3,804.734 W, and the highest value of the estimated heat transfer that occurred during the experiment was 19.535 W, as shown in Table I.7. Refer to Appendix I.2 for the calculations of the actual heat transferred at different gas velocities. The values of the actual heat transferred are tabulated in Table I.7 to I.10. The amount of heat transfer is 80 directly proportional to the mass flow rate (ṁ) and the change in temperature (ΔT) of the volatiles. The ṁ of the volatiles calculated in Appendix A was 2.3 x 10 -3 kg/s while the ṁ in Appendix I was 6.473 x 10 -4 kg/s. The ΔT used in Appendix A was 79°C while the actual ΔT was only about 30°C. The smaller than expected ΔT was due to the heat lost in the gas-exit-pipe of the reactor and in the blower as discussed in Section 4.2.1. Both ṁ and ΔT were higher in the initial calculation which is the reason that the heat transfer was higher. 4.7. Results of Recalculation of the Condenser Length 4.7.1. Comparison of Initial Calculation and Recalculation The recalculated condenser length was much longer than the initial calculation even though the mass flow rate of the volatiles in the recalculation was lower, as discussed in Section 4.6.3. The length of the aluminum condenser in the initial calculation was 78.1 cm while in the recalculation, the length was 264.00 cm. The reason for this is the small value of the convection heat transfer coefficients of the volatiles at the desuperheating and subcooling zones. In the initial calculation the convection coefficient of the volatiles used for the entire condenser was 5,061.875 W/m 2 · K. In the recalculation, the convection coefficients at the desuperheating and subcooling zones were 3.844 W/m 2 · K and 6.495 W/m 2 · K, respectively, and 5,974.601 W/m 2 · K at the condensing zone. Because of the low convection coefficients in the desuperheating and subcooling zones, the condenser must be lengthened to be able to meet the required heat transfer duty. The relationship between the convection coefficient of the volatiles and the cooling water is discussed in Section 4.7.4. The zone with the longest length is the subcooling zone because of the large required ΔT (100°C to 31°C). The length of the subcooling zone is about 94 % of the entire condenser length. Refer to Table H.8 in Appendix H. The experiment results, however, were not comparable with the results of the recalculation because the designed operating temperatures were not realized. The inlet temperature of the volatiles in the experiment was much lower than expected, as discussed in Section 4.2.1. 81 4.7.2. Effect of Flow Velocity If the flow velocity is increased, while keeping all other variables constant, the required condenser length and the pressure drop are increased as shown in Figure 4.17. Small variations in flow velocity have large effect on the required condenser length. The flow velocity seems to be the main variable affecting the required length of the condenser. Therefore, proper blower and piping design is necessary to ensure that condensing system will be able to manage the variations in flow velocity. Figure 4.17: Flow Velocity, Condenser Length, Pressure Drop 4.7.3. Effect of Thermal Conductivity of Condenser Tube The result of the recalculation was different from the initial calculation where the thermal conductivity of the tube material had a significant effect on the condenser length. The thermal conductivity of the tube has negligible effect on the length. If the tube material is 25% Cr & 20% Ni which has a thermal conductivity of 12.80 W/m· K the resulting length is 264.26 cm; if the material is pure silver which has a thermal conductivity of 419 W/m· K the length is 263.99 cm. This is equivalent to a 0.10% change in length. In terms of thermal conductivity, the two materials mentioned are the two extremes shown in Table A-2 of reference [15]. Therefore, there is little significance in using a material with high thermal conductivity. 0 20 40 60 80 100 120 0 200 400 600 800 1,000 1,200 1,400 0 0.25 0.5 0.75 1 C o n d e n s e r L e n g t h , c m Flow Velocity, m/s Length Pressure Drop P r e s s u r e D r o 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 ln 1 1 2 o i i v o w U = d d + + Ah πkL A h ( ) 4.4 where h w represents the convection coefficient of the cooling water. For the sake of illustration, let the 1 st and 2 nd terms of the denominator, and A o assume values of unity. If h w = 1,000, -3 1 = 0.50 1+1+1×10 U ~ 0 1 2 3 4 5 6 260 265 270 275 C o n v e c t i o n C o e f f i c i e n t , 1 0 3 W / m 2 · K Condenser Length, cm 83 That is, for this example only, ( ) lim 1 0.5 ln 1 1 2 o i w i v o w U d d h Ah πkL A h | | | | = = | ÷· | \ . + + For smaller values of h w , however, the overall heat transfer coefficient U decreases. For example, if h w = 10, 1 0.476 1 1 0.1 U = + + 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 (CO 2 ) and Methane (CH 4 ). 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/m 2 · 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 CO 2 and CH 4 . 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 m m t = ( ) A.1 where: %bo = 0.36 t = 23 sec ( )( ) -3 0.36 5 kg 1.304×10 kg s 23sec bo m = = ( ) % pg pg m m t = ( ) A.2 where: %pg = 0.11 t = 23 sec ( )( ) -4 0.11 5 kg = 3.99×10 kg s 23sec pg m = 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,ex Q =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 31 bo Q ÷ ÷ ( ¸ ¸ = 3, 347.680 W bo Q 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,ex Q = ym c T T ÷ ¿ ( ) A.4 ( ) 2 2 pg H pg p,H v,in v,ex Q = y m c T T ÷ where: ṁpg = 3.99x10 -4 kg/s 89 y = 1 2 p,H c = 14,500 J/kg·K Tv,in = 110°C Tv,ex = 31°C ( )( )( )( ) -4 = 1 3.99×10 14, 500 110 31 pg Q ÷ = 457.054 W pg Q The total required heat transfer in the condenser is the sum of Qbo and Qpg, as shown in Eq. (A.5). bo pg Q=Q +Q ( ) A.5 =3,347.680 W+457.054 W Q =3,804.734 W Q 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 1 2 v = gz ( ) A.6 ( )( ) 2 = 2 9.81 1.2 v 2 = 4.852 m s v 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 2 4 ,i π V =v d ( ) A.7 ( ) ( ) 2 = 3.641 0.02096 4 π V -3 3 =1.67×10 m s V 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 Q T = +T ρ Vc ( ) A.8 where: Q = 3,804.734 W ρw = 995.26 kg/m 3 cw = 4,176 J/kg· K Tw,in = 30°C ( )( )( ) -3 3,804.734 = +30 995.26 1.67×10 4,176 w,ex T = 30.546 °C w,ex T The mean temperature difference is computed below in Eq. (A.9). ( ) ( ) 1 2 1 2 Δ Δ Δ Δ ln ln Δ v,in w,ex v,ex w,in lm v,in w,ex v,ex w,in T T T T T T T = = 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 30 Δ = 110 30.546 ln 31 30 lm T ÷ ÷ ÷ ÷ | | | ÷ \ . Δ =17.931°C lm T 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 2 2 2 2 2 = = ,i w annulus i o v d v A v 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 0.0254 w v ÷ = 3.641m s w v The hydraulic diameter of the annulus was solved from Eq. (A.11). The values of Di and do were given above. 2 2 i o H o D - d D = d ( ) A.11 ( ) ( ) 2 2 0.03508 0.0254 = 0.0254 H D ÷ = 0.023 m H D 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 H w w ρ v D = μ ( ) A.12 where: ρw = 995.26 kg/m 3 vw = 3.641 m/s DH = 0.023 m μw = 8.03x10 -4 kg/m·s ( )( )( ) -4 995.26 3.641 0.023 Re = 8.03×10 w Re =104, 018.35 w 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 0.4 Nu = 0.023 Re Pr w w w ( ) A.13 91 ( ) ( ) 0.8 0.4 Nu = 0.023 104, 018.35 5.412 w Nu = 539.671 w Nu w w w H k h = D ( ) A.14 where: k w = 0.619 W/m·°C ( )( ) 539.671 0.619 = 0.023 w h 2 =12,533.625 W m K w h · 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. ( ) 4 Re bo pg i v v v i v m +m d 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 2.3×10 Re = 0.0239 2.82×10 v π Re = 434.501 v ( ) ( ) 1/4 1/4 0.68 sin 0.555 3 fg w sat i w w v w v w i sat i h + c T T ρ ρ ρ g αk h = 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 T Q= = = R R R ÷ ÷ ÷ ( ) A.17 where: 1 v i v R = πd Lh ( ) A.18 ( ) ln 2 o i t t d d R = πk L ( ) A.19 1 w o w R = πd Lh ( ) A.20 The 1 st equation is obtained as follows. v,in i i o v t T T T T = R R ÷ ÷ ( ) ( ) ( ) ( ) 1/4 3 ln 0.68 sin 0.555 2 ÷ ( ÷ ( ÷ ( ( ÷ ¸ ¸ ¸ ¸ 1/ 4 v,in i i o i fg w sat i w w v w i o w i sat i t T T d d d h + c T T ρ ρ ρ g α k T = +T μ d T T k ( ) A.21 92 The 2 nd equation is obtained similarly. i o o w t w T T T T = R R ÷ ÷ ( ) ( ) 2 ln 2 ln t i o w w o i o t o w o i k T d h T + d d T = k d h + d d ( ) A.22 Substituting the 2 nd Equation to the 1 st Equation yields ( ) ( ) ( ) ( ) ( ) ( ) 1/4 1/4 3 2 ln 0.68 sin ln 0.555 2 2 ln t i o w w v,in i i o i fg w sat i w w v w o i i t w i sat i t o w o i k T d h T + T T d d d h + c T T ρ ρ ρ g α k d d T = + k μ d T T k d 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/m 3 hfg 2,257,000 J/kg ρv 0.5506 kg/m 3 hw 12,533.625 W/m 2 -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/s 2 cw 4,195 J/kg-K Applying Bisection Method to Eq. (A.23) to solve for Ti yields = 49.931°C i T Ti is substituted back to Eq. (A.16) to solve for the volatile convection heat transfer coefficient. 2 = 5, 061.875 W m K v h · 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. ( ) ln 1 1 Δ 2 o i lm i v t o w d d Q L= + + π T d h k d h ( ( ¸ ¸ ( ) A.24 where: Q = 3,804.734 W ΔTlm = 17.931°C hw = 12,533.625 W/m 2 ·°C hv = 5,061.875 W/m 2 ·°C kt = 204 W/m·°C do = 0.0254 m di = 0.0239 m ( ) ( )( ) ( ) ( ) ( )( ) ln 0.0254 0.0239 3,804.734 1 1 = + + 17.931 0.0239 5, 061.875 2 204 0.0254 12, 533.625 L π ( ( ( ¸ ¸ = 0.781m= 78.1cm L The same procedures above were followed in calculating the length of the stainless steel condenser. The length of the stainless condenser was determined to be 0.999 m or 99.9 cm. 93 Appendix B. Fabricated Condenser Parts and Assembly B.1. Dimensions and Parts Figure B.1: Aluminum Condenser Figure B.2: Stainless Condenser Note: All dimensions above are in cm Table B.1: List of Parts Part # Part Name Material Size 1 Outer Tube GI pipe 1-1/4 in. 2 Inner Tube Aluminum/Stainless pipe 1 in. 3 Cross Tee GI cross tee 1-1/4 in. 4 Adapter Nipple GI nipple 1 x 1 in. 5 Water Inlet/Exit Nipples GI nipple ¾ x 2 in. 6 Bushing Reducer GI bushing reducer 1-1/4 x 1 in. 94 Figure B.3: Exploded View of Condenser Table B.2: Inner-Tube Actual Dimensions Aluminum (1 in.) Stainless (1 in.) Inner Dia. (di), m Outer Dia. (do), m Inner Dia. (Di), m Outer Dia. (Do), m 0.0239 0.0254 0.0226 0.0254 Table B.3: Outer-Tube Actual Dimensions Outer tube (1-1/4 in.) Water Inlet/Exit Tube (3/4 in.) Inner Dia. (Di), m Outer Dia. (Do), m Inner Dia. (d2,i), m 0.03508 0.0425 0.02096 B.2. Condenser Accessories The adapter is an accessory of the condenser where the bio-oil and pyrolysis gas are isolated. Figure B.4 shows the adapter. By density difference the heavier liquid bio-oil flows below the lighter pyrolysis gas. When they reach the adapter, the bio-oil drops down, as indicated in Figure B.4, to the beaker. When the oil-valve is closed the pyrolysis gas flow directly out through the gas-exit- valve. The inner diameter of the gas-exit-valve is 7.5 mm. Figure B.4: Adapter 95 The static pressure tap was another accessory of the condenser which was used in measuring the static pressure of the flow. The manometer tube was attached to the valve indicated in Figure B.5. Figure B.5: Static Pressure Tap B.3. Thermocouple Probes and Pressure Taps Figure B.6: Position of Thermocouple Probes Figure B.7: Position of Static Pressure Taps a) Aluminum Condenser b) Stainless Condenser 96 B.4. Condenser Tilt Angle The actual tilt angle was measure by using a water-hose level gage similar to Figure 3.13 and a straight edge, in this case, a triangle. The two points in the water-hose level gage where the top of the water column rest represent two points on the horizontal. These two points were connected by the triangle. The triangle then represents a horizontal line. A photograph of this set-up, shown in Figure B.8, was taken with the camera positioned at approximately the same elevation as the set-up. In the photograph, lines parallel to the triangle and the condenser were drawn. The angle between these lines, which was measured with a protractor, is the tilt angle of the condenser. Figure B.8: Condenser Tilt Angle 97 Appendix C. Cooling Water Flow Rate Measurements Table C.1: Mass Flow Rate for Fully Open Trial Mass, kg Time, sec Mass Flow, kg/s 1 5.734 11.74 0.488 2 5.432 11.37 0.478 3 4.984 10.4 0.479 4 5.374 11.64 0.462 5 5.058 11.63 0.435 Average 0.468 Table C.2: Mass Flow Rate for One Valve-Turn Trial Mass, kg Time, sec Mass Flow, kg/s 1 5.364 11.4 0.471 2 4.936 10.58 0.467 3 4.954 10.13 0.489 4 4.842 10.3 0.470 5 4.852 10.33 0.470 Average 0.473 Table C.3: Mass Flow Rate for Two Valve-Turns Trial Mass, kg Time, sec Mass Flow, kg/s 1 4.252 10.43 0.408 2 4.166 10.35 0.403 3 4.2 10.26 0.409 4 4.26 10.35 0.412 5 4.238 10.41 0.407 Average 0.408 Table C.4: Mass Flow Rate for Three Valve-Turns Trial Mass, kg Time, sec Mass Flow, kg/s 1 2.99 10.06 0.297 2 3.146 10.46 0.301 3 3.094 10.29 0.301 4 3.136 10.37 0.302 5 3.072 10.35 0.297 Average 0.300 Table C.5: Mass Flow Rate for Four Valve-Turns Trial Mass, kg Time, sec Mass Flow, kg/s 1 1.59 10.35 0.154 2 1.48 10.17 0.146 3 1.618 10.92 0.148 4 1.524 10.36 0.147 5 1.538 10.25 0.150 Average 0.149 98 Appendix D. Static Pressure and Gas Velocity Measurements D.1. Static Pressure Measurements Table D.1: Manometer Reading for Run A4 (Gas-Exit-Valve Full Open) Manometer Tube Manometer Reading, inches H2O Average, inches H2O Pt. 1 Pt. 2 Upper 1-1/2 1-14/16 1.6875 Lower 2-2/16 2-1/2 2.3125 Table D.2: Manometer Reading for Run A4 (Gas-Exit-Valve Slightly Close) Manometer Tube Manometer Reading, inches H2O Average, inches H2O Pt. 1 Pt. 2 Upper 10/16 1 0.8125 Lower 2-14/16 3-4/16 3.0625 Table D.3: Manometer Reading for Run A5 (Gas-Exit-Valve Full Open) Manometer Tube Manometer Reading, inches H2O Average, inches H2O Pt. 1 Pt. 2 Upper 1-15/16 2-5/16 2.12375 Lower 1-10/16 2 1.8125 Table D.4: Manometer Reading for Run A5 (Gas-Exit-Valve Slightly Close) Manometer Tube Manometer Reading, inches H2O Average, inches H2O Pt. 1 Pt. 2 Upper 1 1-6/16 1.1875 Lower 2-8/16 2-14/16 2.6875 Since the manometer was inclined at approximately 30° from the horizontal, Eq. (D.1) was used to solve the actual height of the water column. Δ sin 30° t t h u l ÷ = ( ) D.1 Δ 1.6875 2.3125 sin 30° h = ÷ 2 Δ 0.3125 in. H O h = Converting the height of the water column to, 2 101325 Pa Δ 408 in.H O s p h | | = | \ . 2 2 101325 Pa 0.3125 in. H O 408 in.H O s p | | = | \ . 77.608 Pa s p = 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 ex Av A v = ( ) D.2 2 2 nozzle ex nozzle ex i i i A v d v v A d = = ( ) ( ) ( ) 2 -3 2 7.5×10 m 1, 700 ft min 1m 1min 3.28 ft 60 sec 0.0239 m i v | || | = | | \ .\ . 0.851m s i v = Table D.6: Velocity Inside Inner-tube Run No. Gas Exit Velocity, m/s Fully Open Slightly Close A4 0.851 0.325 A5 0.500 0.125 100 Appendix E. Pyrolysis Products E.1. Periodic Bio-oil Volume Measurement Table E.1: Bio-oil Volume Collected for Run A1 Run Duration, h:mm Volume, ml Run Duration, h:mm Volume, ml 0:42 Start 3:12 55 0:57 34 3:27 42 1:12 33 3:42 30 1:27 31 3:57 41 1:42 26 4:12 49 1:57 12 4:27 47 2:12 20 4:42 36 2:27 16 4:57 32 2:42 42 5:12 28 2:57 44 Total: 618 ml Table E.2: Bio-oil Volume Collected for Run A2 Run Duration, h:mm Volume, ml Run Duration, h:mm Volume, ml 0:10 Sart 2:25 34 0:25 5 2:40 22 0:40 3 2:55 35 0:55 5 3:10 35 1:10 8 3:25 32 1:25 15 3:40 34 1:40 14 3:55 48 1:55 38 4:10 16 2:10 23 Total: 367 ml Table E.3: Bio-oil Volume Collected for Run A3 Run Duration, h:mm Volume, ml Run Duration, h:mm Volume, ml 0:07 Start 1:22 17 0:22 26 1:37 60 0:37 20 1:52 22 0:52 17 2:07 36 1:07 18 Table E.4: Bio-oil Volume Collected for Run A4 Run Duration, h:mm Volume, ml Run Duration, h:mm Volume, ml 1:39 Start 2:39 70 1:54 29 2:54 30 2:09 39 3:09 35 2:24 33 3:24 14 Leak: 4 ml Total: 254 ml Table E.5: Bio-oil Volume Collected for Run A5 Run Duration, h:mm Volume, ml Run Duration, h:mm Volume, ml 0:11 Start 3:11 12 0:26 31 3:26 63 0:41 33 3:41 38 0:56 21 3:56 36 1:11 12 4:11 21 1:26 12 4:26 13 1:41 16 4:41 18 1:56 28 4:56 21 2:11 13 5:11 27 2:26 20 5:26 21 2:41 33 5:41 39 2:56 13 5:56 38 Total: 579 ml Table E.6: Bio-oil Volume Collected for Run A6 Run Duration, h:mm Volume, ml Run Duration, h:mm Volume, ml 1:56 0 4:26 26 2:11 15 4:41 44 2:26 25 4:56 27 2:41 24 5:11 26 2:56 28 5:26 31 3:11 37 5:41 24 3:26 45 5:56 32 3:41 35 6:11 30 3:56 42 6:26 29 4:11 28 6:41 24 Leak: 17 ml Total: 589 ml 101 Table E.7: Bio-oil Volume Collected for Run A7 Run Duration, h:mm Volume, ml Run Duration, h:mm Volume, ml 1:08 Start 3:23 44 1:23 2 3:38 35 1:38 4 3:53 35 1:53 25 4:08 9 2:08 18 4:23 20 2:23 16 4:38 15 2:38 36 4:53 50 2:53 30 5:08 28 3:08 39 5:23 38 Leak: 24 ml Total: 468 ml Table E.8: Bio-oil Volume Collected for Run A8 Run Duration, h:mm Volume, ml Leakage, ml 0:20 Start 0:35 24 52 0:50 29 51 1:05 11 21 1:20 29 25 1:35 6 18 1:50 24 23 2:05 11 10 2:20 58 18 2:35 41 9 2:50 28 4 3:05 18 4 3:20 40 4 3:35 26 6 3:50 10 4:05 13 4:20 32 Leak: 245 ml Total: 645 ml Table E.9: Bio-oil Volume Collected for Run S5 Run Duration, h:mm Volume, ml Run Duration, h:mm Volume, ml 0:53 Start 3:38 49 1:08 28 3:53 41 1:23 29 4:08 18 1:38 29 4:23 26 1:53 29 4:38 23 2:08 33 4:53 55 2:23 25 5:08 57 2:38 24 5:23 35 2:53 24 5:38 34 3:08 15 5:48 28 3:23 23 Total: 625 ml Table E.10: Bio-oil Volume Collected for Run S6 Run Duration, h:mm Volume, ml Run Duration, h:mm Volume, ml 0:08 Start 3:08 72 0:23 75 3:23 39 0:38 29 3:38 23 0:53 18 3:53 16 1:08 13 4:08 58 1:23 16 4:23 46 1:38 18 4:38 30 1:53 10 4:53 34 2:08 16 5:08 39 2:23 13 5:23 58 2:38 8 5:38 58 2:53 12 5:53 55 Total: 756 ml Table E.11: Bio-oil Volume Collected for Run S7 Run Duration, h:mm Volume, ml Run Duration, h:mm Volume, ml 2:17 Start 3:47 66 2:32 11 4:02 51 2:47 11 4:17 40 3:02 28 4:32 45 3:17 27 4:43 12 3:32 31 Total: 322 ml 102 E.2. Bio-oil and Pyrolysis Gas Yield Table E.12: Bio-oil Volume Collected for Run S8 Run Duration, h:mm Volume, ml Run Duration, h:mm Volume, ml 0:06 Start 2:36 10 0:21 46 2:51 19 0:36 36 3:06 15 0:51 22 3:21 40 1:06 40 3:36 44 1:21 27 3:51 52 1:36 23 4:06 48 1:51 13 4:21 30 2:06 12 4:36 48 2:21 25 Total: 550 ml Table E.13: Mass of Marine Florae Feedstock and Pyrolysis Products Run # Type of Feedstock Weight, kg Feedstock Char Oil Gas A1 Pure Green Pellets 5.000 2.172 0.600 2.228 A2 Pure Red Pellet 4.474 3.639 0.346 0.489 A4 Pure Green Pellets 5.000 3.768 0.228 1.004 A5 Red Raw 5.000 3.334 0.564 1.102 A6 Pure Seagrass Pellets 4.968 3.174 0.588 1.206 A7 Pure Brown Pellets 5.000 3.700 0.452 0.848 A8 Seagrass with Binder 5.000 3.406 0.638 0.956 S1 Brown with Binder 5.000 3.23 0.652 1.118 S2 Seagrass with Binder 5.000 2.546 0.984 1.470 S3 Red Raw 3.918 2.954 0.408 0.556 S4 Pure Red Pellets 3.776 2.949 0.554 0.273 S5 Pure Green Pellets 5.095 3.984 0.606 0.505 S6 Green Raw 4.998 3.296 0.724 0.978 S7 Pure Brown Pellets 4.858 3.570 0.310 0.978 S8 Pure Seagrass Pellets 5.000 3.410 0.530 1.060 Table E.14: Mass Percentage of Pyrolysis Products Run # Type of Feedstock Product Percentage, % Char Oil Gas A1 Pure Green Pellets 43.44 12.00 44.56 A2 Pure Red Pellet 81.34 7.73 10.93 A4 Pure Green Pellets 75.36 4.56 20.08 A5 Red Raw 66.68 11.28 22.04 A6 Pure Seagrass Pellets 63.89 11.84 24.28 A7 Pure Brown Pellets 74.00 9.04 16.96 A8 Seagrass with Binder 68.12 12.76 19.12 S1 Brown with Binder 64.60 13.04 22.36 S2 Seagrass with Binder 50.92 19.68 29.40 S3 Red Raw 75.40 10.41 14.19 S4 Pure Red Pellets 78.10 14.67 7.23 S5 Pure Green Pellets 78.19 11.89 9.91 S6 Green Raw 65.95 14.49 19.57 S7 Pure Brown Pellets 73.49 6.38 20.13 S8 Pure Seagrass Pellets 68.20 10.60 21.20 Average 68.51 11.36 20.13 103 Table E.15: Density of Bio-oil Run # Type of Feedstock Density, kg/m3 A1 Pure Green Pellets 970.87 A2 Pure Red Pellet 942.78 A4 Pure Green Pellets 897.64 A5 Red Raw 974.09 A6 Pure Seagrass Pellets 998.30 A7 Pure Brown Pellets 965.81 A8 Seagrass with Binder 989.15 S1 Brown with Binder 1018.75 S2 Seagrass with Binder 1004.08 S3 Red Raw 1020.00 S4 Pure Red Pellets 1045.28 S5 Pure Green Pellets 969.60 S6 Green Raw 957.67 S7 Pure Brown Pellets 962.73 S8 Pure Seagrass Pellets 963.64 Average 978.69 E.3. Product Composition and Residence Time from Añora (2010) Table E.16: Mass Percentage and Residence Time for Green Algae [1] Green 80/20 Pellets Pure Green Pellets Trial 1st 2nd 3rd 1st 2nd 3rd %bo 0.32 0.32 0.32 0.4 0.4 0.4 %pg 0.16 0.16 0.16 0.08 0.08 0.08 t, min 33 32 30 28 27 29 Table E.17: Mass Percentage and Residence Time for Red Algae [1] Red 80/20 Pellets Pure Red Pellets Trial 1st 2nd 3rd 1st 2nd 3rd %bo 0.15 0.15 0.15 0.24 0.24 0.24 %pg 0.32 0.32 0.32 0.21 0.21 0.21 t, min 17 23 25 26 23 25 Table E.18: Mass Percentage and Residence Time for Brown Algae [1] Brown 80/20 Pellets Pure Brown Pellets Trial 1st 2nd 3rd 1st 2nd 3rd %bo 0.36 0.36 0.36 0.36 0.36 0.36 %pg 0.11 0.11 0.11 0.11 0.11 0.11 t, min 29 25 26 55 23 28 Table E.19: Mass Percentage and Residence Time for Seagrass [1] Seagrass 70/30 Pellets Pure Seagrass Pellets Trial 1st 2nd 3rd 1st 2nd 3rd %bo 0.32 0.32 0.32 0.28 0.28 0.28 %pg 0.19 0.19 0.19 0.23 0.23 0.23 t, min 61 28 33 47 29 30 104 Appendix F. Volatile Temperature Graph F.1. Volatile Temperature Graph with Plotted Periodic Bio-oil Yield F i g u r e F . 1 : V o l a t i l e T e m p e r a t u r e G r a p h o f R u n A 1 105 F i g u r e F . 2 : V o l a t i l e T e m p e r a t u r e G r a p h o f R u n A 2 106 F i g u r e F . 3 : V o l a t i l e T e m p e r a t u r e G r a p h o f R u n A 3 F i g u r e F . 4 : V o l a t i l e T e m p e r a t u r e G r a p h o f R u n A 4 107 F i g u r e F . 5 : V o l a t i l e T e m p e r a t u r e G r a p h o f R u n A 5 108 F i g u r e F . 6 : V o l a t i l e T e m p e r a t u r e G r a p h o f R u n A 6 109 F i g u r e A . 7 : V o l a t i l e T e m p e r a t u r e G r a p h o f R u n A 7 110 F i g u r e F . 8 : V o l a t i l e T e m p e r a t u r e G r a p h o f R u n A 8 111 F i g u r e F . 9 : V o l a t i l e T e m p e r a t u r e G r a p h o f R u n S 5 112 F i g u r e F . 1 0 : V o l a t i l e T e m p e r a t u r e G r a p h o f R u n S 6 113 F i g u r e F . 1 1 : V o l a t i l e T e m p e r a t u r e G r a p h o f R u n S 7 114 F i g u r e F . 1 2 : V o l a t i l e T e m p e r a t u r e G r a p h o f R u n S 8 115 F.2. Volatile Temperature Graph without Plotted Periodic Bio-oil Yield F i g u r e F . 1 3 : V o l a t i l e T e m p e r a t u r e G r a p h o f R u n S 1 116 F i g u r e F . 1 4 : V o l a t i l e T e m p e r a t u r e G r a p h o f R u n S 2 117 F i g u r e F . 1 5 : V o l a t i l e T e m p e r a t u r e G r a p h o f R u n S 3 118 F i g u r e F . 1 6 : V o l a t i l e T e m p e r a t u r e G r a p h o f R u n 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 s TP G L v 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. 1 1 1 p RT µ = ( ) G.2 2 2 2 p RT µ = ( ) 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, 77.608 1.650 kg m 188.9 51.5 273 CO µ +101, 058.355 = = + ( )( ) 2 3 2, 77.608 1.762 kg m 188.9 30.9 273 CO µ +101, 058.355 = = + For CH4, R is 518.2 J/kg·°C. The densities are ( )( ) 4 3 1, 77.608 0.601kg m 518.2 51.5 273 CH µ +101, 058.355 = = + ( )( ) 4 3 2, 77.608 0.642 kg m 518.2 30.9 273 CH µ +101, 058.355 = = + Run A4 had Pure Green Pellet feedstock which has a composition of 92.08% CO2 and 7.92% CH4. Refer to Table 4.6. The density of the gas mixture was calculated using Eq. (G.4). 2 2 4 4 G CO CO CH CH y y µ µ µ = + ( ) G.4 where: 2 CO y = 0.9208 4 CH y = 0.0792 At the inlet, ( )( ) ( )( ) 3 1, 0.9208 1.650 0.0792 0.601 1.567 kg m G µ = + = and at the exit, ( )( ) ( )( ) 3 2, 0.9208 1.762 0.0792 0.642 1.673kg m G µ = + = 120 The average pyrolysis gas density, therefore, is 3 1.567 1.673 1.620 kg m 2 G µ + = = The absolute viscosities of each gas component are obtained from gas property tables by linear interpolation. Table G.1 shows the absolute viscosities of CO2 and CH4 at 51.5°C and 30.9°C. Table G.1: Absolute Viscosities of Pyrolysis Gas Components Gas Absolute Viscosity, kg/m·s 51.5°C 30.9°C CO2 1.606 x 10 -5 1.513 x 10 -5 CH4 1.189 x 10 -5 1.131 x 10 -5 The absolute viscosity of the gas mixture is calculated using Eq. (G.5). 2 2 4 4 G CO CO CH CH y y µ µ µ = + ( ) G.5 At the inlet, ( )( ) ( )( ) 5 5 5 1, 0.9208 1.606 10 0.0792 1.189 10 1.573 10 kg m s G µ ÷ ÷ ÷ = × + × = × · and at the exit, ( )( ) ( )( ) 5 5 5 2, 0.9208 1.513 10 0.0792 1.131 10 1.483 10 kg m s G µ ÷ ÷ ÷ = × + × = × · The average pyrolysis gas viscosity, therefore, is 5 5 5 1.573 10 1.483 10 1.528 10 kg m s 2 G µ ÷ ÷ ÷ × + × = = × · 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 bo L V V = ( ) G.6 where: Vbo = 15 ml 3 8 3 29 ml 1 1L 1m 3.222 10 m s 15 min 60 s 1000 ml 1000 L L V ÷ | | | || | = = × | | | \ .\ .\ . The flow area occupied by the bio-oil is L bo TP V A v = ( ) G.7 where: vTP = 0.851 m/s 8 8 2 3.222 10 3.786 10 m 0.851 bo A ÷ ÷ × = = × The flow area of the pyrolysis gas is 2 4 pg i bo i bo A A A d A t = ÷ = ÷ ( ) G.8 where: di = 0.0239 m 121 ( ) 2 8 4 2 0.0239 3.786 10 4.486 10 m 4 pg A t ÷ ÷ = ÷ × = × The volume flow rate of the pyrolysis gas is calculated using Eq. (G.9). G pg TP V A v = ( ) G.9 ( )( ) 4 4 3 4.486 10 0.851 3.818 10 m s G V ÷ ÷ = × = × The void fraction is calculated using Eq. (G.10). G G L G V V V c = + ( ) G.10 4 8 4 3.818 10 0.99992 3.222 10 3.818 10 G c ÷ ÷ ÷ × = = × + × The quality is calculated using Eq. (G.11). ρL is the actual density of Pure Green Pellets in Run A4. 1 G G L G G G L x µ c µ µ c c µ | | | \ . = | | ÷ + | \ . ( ) G.11 where: ρL = 897.64 kg/m 3 ρG = 1.620 kg/m 3 εG = 0.99992 1.620 0.99992 897.64 0.955 1.620 1 0.99992 0.99992 897.64 x | | | \ . = = | | ÷ + | \ . The density of the two-phase mixture is calculated from Eq. (G.12). ( ) 1 G L TP L G x x µ µ µ µ µ = + ÷ ( ) G.12 where: ρL = 897.64 kg/m 3 ρG = 1.620 kg/m 3 x = 0.955 ( )( ) ( )( ) ( )( ) 3 1.620 897.64 1.696 kg m 0.955 897.64 1 0.955 1.620 TP µ = = + ÷ Similarly, the absolute viscosity is calculated from Eq. (G.13). The value of μL is evaluated at 41.2°C for water. ( ) 1 G L TP L G x x µ µ µ µ µ = + ÷ ( ) 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 10 1.598 10 kg m s 0.955 6.413 10 1 0.955 1.528 10 TP µ ÷ ÷ ÷ ÷ ÷ × × = = × · × + ÷ × 122 The mass flux of is calculated from Eq. (G.14), where A is the cross sectional area of the inner tube. L G TP TP V V m A µ | | + = | \ . - ( ) G.14 where: L V = 3.222 x 10 -8 m 3 /s G V = 3.818 x 10 -4 m 3 /s ρTP = 1.696 kg/m 3 A = 4.486 x 10 -4 m 2 ( ) 8 4 2 4 3.222 10 3.818 10 1.696 1.443 kg m s 4.486 10 TP m ÷ ÷ ÷ | | × + × = = · | × \ . - The Reynolds number is calculated using Eq. (G.15). Re TP i TP TP m d µ = - ( ) G.15 where: TP m - = 1.443 kg/m 2 ·s μTP = 1.598 x 10 -5 kg/m·s di = 0.0239 m ( )( ) ( ) 5 1.443 0.0239 Re 2,158.809 1.598 10 TP ÷ = = × For two-phase flow, Re > 2,000 is turbulent. [14] Thus, the friction factor is calculated using Eq. (G.16). 1/4 0.079Re TP TP f ÷ = ( ) G.16 ( ) 1/4 0.079 2,158.809 0.0116 TP f ÷ = = 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 2 sin TP TP TP i TP f m L p g L d µ o µ A = ÷ - ( ) G.17 where: fTP = 0.0116 TP m - = 1.443 kg/m 2 ·s L = 0.738 m di = 0.0239 m ρTP = 1.696 kg/m 3 α = -25° g = 9.81 m/s 2 ( )( ) ( ) ( )( ) ( )( )( ) ( ) 2 2 0.0116 1.443 0.738 9.81 1.696 0.738 sin 25 0.0239 1.696 p A = ÷ ÷ ° 6.067 Pa p A = 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 1 p p p = ÷A ( ) G.18 where: p1 = 77.608 Pa (gage) 123 2 77.608 6.067 71.541Pa (gage) p = ÷ = p2 is inserted back to Eq. (G.3) to recalculate ρ2. The first iteration yields a 5.977 x 10 -3 % change in the ρ2. This means that the first solution was already a sufficient estimate of the pressure drop and the mass flux which is used to calculate the actual heat transfer. For Run A4 and „full open‟ gas-exit-valve, the time when the static pressure and gas velocity were measured is from 1:43:00 to 1:44:50. In between this time duration there are 12 temperature readings at a 10-second interval. Since the gas properties are highly sensitive to temperature, 12 pressure drop calculations were done corresponding to each of the 12 temperature readings. The results of these calculations are presented in Table G.2. The same calculations were also conducted for Run A4 „slightly closed‟ and Run A5 both „full open‟ and „slightly closed‟. The summary is shown in Table G.3, G.4, and G.5. Table G.2: Summary of Pressure Drop for Run A4, „full open‟ Time, h:min:sec Gas Temperature, °C Pressure Drop, Pa Inlet Exit 1:43:00 51.5 30.9 6.067 1:43:10 49.6 28.5 6.104 1:43:20 48.6 28.1 6.116 1:43:30 47.8 28 6.124 1:43:40 47.3 27.9 6.129 1:43:50 46.6 27.9 6.134 1:44:00 46.2 27.8 6.139 1:44:10 45.8 27.8 6.142 1:44:20 60.2 51.3 5.551 1:44:30 65.3 55.1 5.495 1:44:40 65.9 40.5 5.871 1:44:50 60.1 30 6.008 Average 5.990 Table G.3: Summary of Pressure Drop for Run A4, „slightly closed‟ Time, h:min:sec Gas Temperature, °C Pressure Drop, Pa Inlet Exit 1:45:00 57.6 28.7 5.773 1:45:10 56.3 28.4 5.784 1:45:20 49.2 28.3 5.835 1:45:30 48.3 28.2 5.842 1:45:40 60 31.9 5.731 1:45:50 60.9 32.5 5.720 1:46:00 61.8 33 5.710 1:46:10 63.6 33.7 5.693 1:46:20 64.2 30.9 5.711 1:46:30 64.1 30.1 5.719 1:46:40 60.3 29.3 5.750 1:46:50 57.7 28.3 5.776 Average 5.754 124 G.2. Actual Heat Transfer In the calculation of the actual heat transferred, the specific heats at constant pressure of the individual gas components were determined using Eq. (G.19). [11] 2 3 a bT cT dT c 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 3 2 5 9 1000 22.26 5.981 10 314.2 3.501 10 314.2 7.649 10 314.2 44.010 CO c ÷ ÷ ÷ ( = + × + ÷ × + × ¸ ¸ 2 859.526 J kg K CO c = · ( )( ) ( )( ) ( )( ) 4 2 3 2 5 10 1000 19.89 5.024 10 314.2 1.269 10 314.2 1.101 10 314.2 16.043 CH c ÷ ÷ ÷ ( = + × + × + ÷ × ¸ ¸ 4 2,301.613J kg K CH c = · The specific heat of the gas mixture is calculated using Eq. (G.20) 2 2 4 4 G CO CO CH CH c y c y c = + ( ) G.20 where: 2 CO y = 0.9208 4 CH y = 0.0792 Table G.4: Summary of Pressure Drop for Run A5, „full open‟ Time, h:min:sec Gas Temperature, °C Pressure Drop, Pa Inlet Exit 0:15:30 58.8 38.5 5.038 0:15:40 61.8 40.7 5.052 0:15:50 63.6 41.3 5.038 0:16:00 64.6 42 5.028 0:16:10 65.8 42.7 5.016 0:16:20 67.2 43.7 5.002 0:16:30 68.4 44.5 4.991 0:16:40 69.2 45.2 4.982 0:16:50 70 45.6 4.975 0:17:00 70.1 34.9 5.044 Average 5.017 Table G.5: Summary of Pressure Drop for Run A5, „slightly closed‟ Time, h:min:sec Gas Temperature, °C Pressure Drop, Pa Inlet Exit 0:18:00 67.4 29.4 6.177 0:18:10 67.6 29.9 6.234 0:18:20 67.9 30.7 6.227 0:18:30 68.6 31.4 6.218 0:18:40 69 31.7 6.213 0:18:50 69.4 31.7 6.211 0:19:00 69.6 31.8 6.209 0:19:10 69.8 31.7 6.208 0:19:20 70.2 43.7 6.123 0:19:30 70.3 34.4 6.186 Average 6.201 125 ( )( ) ( )( ) 0.9208 859.526 0.0792 2,301.613 973.709 J kg K G c = + = · 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 L TP L G c c c xc x c = + ÷ ( ) G.21 where: cG = 973.709 J/kg·K cL = 4,181.289 J/kg·K x = 0.955 ( )( ) ( )( ) ( )( ) 973.709 4,181.289 1, 008.266 J kg K 0.955 4,181.289 1 0.955 973.709 TP c = = · + ÷ 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 TP Q m Ac T = A - ( ) G.22 where: TP m - = 1.443 kg/m 2 ·s A = 4.486 x 10 -4 m 2 cTP = 1,008.266 J/kg·K ΔT = 51.5°C - 30.9°C = 20.6°C ( )( )( )( ) 4 1.443 4.486 10 1, 008.266 20.6 13.446 W Q ÷ = × = The heat transfer for Runs A4 and A5 during the time when the static pressure readings were taken is summarized in Table G.7, G.8, G.9, and G.10. Table G.7: Summary of Heat Transfer for Run A4, „full open‟ Time, h:min:sec Gas Temperature, °C Heat Transfer, W Inlet Exit 1:43:00 51.5 30.9 13.446 1:43:10 49.6 28.5 13.827 1:43:20 48.6 28.1 13.450 1:43:30 47.8 28 13.001 1:43:40 47.3 27.9 12.744 1:43:50 46.6 27.9 12.292 1:44:00 46.2 27.8 12.100 1:44:10 45.8 27.8 11.841 1:44:20 60.2 51.3 5.660 1:44:30 65.3 55.1 6.439 1:44:40 65.9 40.5 16.243 1:44:50 60.1 30 19.535 Average 12.548 Table G.8: Summary of Heat Transfer for Run A4, „slightly closed‟ Time, h:min:sec Gas Temperature, °C Heat Transfer, W Inlet Exit 1:45:00 57.6 28.7 8.133 1:45:10 56.3 28.4 7.859 1:45:20 49.2 28.3 5.914 1:45:30 48.3 28.2 5.692 1:45:40 60 31.9 7.875 1:45:50 60.9 32.5 7.951 1:46:00 61.8 33 8.055 1:46:10 63.6 33.7 8.349 1:46:20 64.2 30.9 9.317 1:46:30 64.1 30.1 9.520 1:46:40 60.3 29.3 8.705 1:46:50 57.7 28.3 8.276 Average 7.970 126 Table G.9: Summary of Heat Transfer for Run A5, „full open‟ Time, h:min:sec Gas Temperature, °C Heat Transfer, W Inlet Exit 0:15:30 58.8 38.5 7.414 0:15:40 61.8 40.7 7.752 0:15:50 63.6 41.3 8.179 0:16:00 64.6 42 8.279 0:16:10 65.8 42.7 8.451 0:16:20 67.2 43.7 8.583 0:16:30 68.4 44.5 8.717 0:16:40 69.2 45.2 8.744 0:16:50 70 45.6 8.883 0:17:00 70.1 34.9 12.929 Average 8.793 Table G.10: Summary of Heat Transfer for Run A5, „slightly closed‟ Time, h:min:sec Gas Temperature, °C Heat Transfer, W Inlet Exit 0:18:00 67.4 29.4 5.396 0:18:10 67.6 29.9 5.392 0:18:20 67.9 30.7 5.319 0:18:30 68.6 31.4 5.317 0:18:40 69 31.7 5.330 0:18:50 69.4 31.7 5.387 0:19:00 69.6 31.8 5.401 0:19:10 69.8 31.7 5.444 0:19:20 70.2 43.7 3.772 0:19:30 70.3 34.4 5.124 Average 5.188 127 Appendix H. Recalculation of Double-Pipe Condenser Length H.1. Bio-oil and Pyrolysis Gas Properties The calculations presented here are for Seagrass with Binder which has a composition of 90.24% CO2 and 9.76% CH4. First, the properties of the bio-oil and pyrolysis gas were determined. The properties of the bio-oil were obtained from property tables of water and listed below according to the condenser zone in which they were applied. In the desuperheating zone the properties were evaluated at 110°C and 100°C but only the average values are shown in Table H.1, except for the enthalpies. Table H.2 shows the bio- oil properties applied in the condensing zone and Table H.3 shows the average bio-oil properties applied in the subcooling zone. Table H.1: Bio-oil Properties Applied in Desuperheating Zone Property Symbol Numerical Value Unit Enthalpy at 110°C hi 2,696,200 J/kg Enthalpy at 100°C hsat 2,676,100 J/kg Absolute Viscosity μL,des 1.244 x 10 -5 kg/m·s Density ρL,des 0.5816 kg/m 3 Thermal Conductivity kL,des 0.02519 W/m·K Table H.2: Bio-oil Properties Applied in Condensing Zone Property Symbol Numerical Value Unit Viscosity of Saturated Steam μv,con 1.227 x 10 -5 kg/m·s Viscosity of Saturated Water μL,con 2.826 x 10 -4 kg/m·s Thermal Conductivity of Saturated Water kL,con 0.682 W/m·K Density of Saturated Steam ρL,con 0.590 kg/m 3 Density of Saturated Water ρv,con 958.320 kg/m 3 Latent Heat of Vaporization hfg 2,257,000 J/kg Specific Heat of saturated Water cL,con 4,211 J/kg·K Table H.3: Bio-oil Properties Applied in Subcooling Zone Property Symbol Numerical Value Unit Absolute Viscosity μL,sub 5.342 x 10 -4 kg/m·s Density ρL,sub 976.709 kg/m 3 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 3 a bT cT dT c 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 3 2 5 9 1000 22.26 5.981 10 383 3.501 10 383 7.649 10 383 44.010 CO c ÷ ÷ ÷ ( = + × + ÷ × + × ¸ ¸ 2 919.138 J kg K CO c = · 128 ( )( ) ( )( ) ( )( ) 4 2 3 2 5 10 1000 19.89 5.024 10 383 1.269 10 383 1.101 10 383 16.043 CH c ÷ ÷ ÷ ( = + × + × + ÷ × ¸ ¸ 4 2,554.835 J kg K CH c = · The specific heat of the gas mixture was calculated using Eq. (H.2). The values of the mass fraction are for Seagrass with Binder shown in Table 4.6. 2 2 4 4 G CO CO CH CH c y c y c = + ( ) H.2 where: 2 CO y = 0.9024 4 CH y = 0.0976 2 CO c = 919.138 J/kg·K 4 CH c = 2,554.835 J/kg·K ( )( ) ( )( ) ,110 0.9024 919.138 0.0976 2,554.835 1, 078.788 J kg K G C c ° = + = · 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 K CO c = · 4 2,517.569 J kg K CH c = · ,100 1, 067.658 J kg K G C c ° = · 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, 067.658 1, 073.223 J kg K 2 G des c + = = · 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 µ = ( ) 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 3 101, 511.259 1.403 kg m 188.9 383 CO µ = = ( )( ) 4 3 101, 511.259 0.511kg m 518.2 383 CH µ = = The density of the pyrolysis gas was calculated using Eq. (H.4). 2 2 4 4 G CO CO CH CH y y µ µ µ = + ( ) H.4 where: 2 CO y = 0.9024 4 CH y = 0.0976 2 CO µ = 1.403 kg/m 3 4 CH µ = 0.511 kg/m 3 129 ( )( ) ( )( ) 3 ,110 0.9024 1.403 0.0976 0.511 1.316 kg m G C µ ° = + = The density of the gas mixture at 100°C was also calculated using Eq. (H.3) and (H.4). The results are 2 3 1.441kg m CO µ = 4 3 0.525 kg m CH µ = 3 ,100 1.351kg m G C µ ° = 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 1.351 1.334 kg m 2 G des µ + = = The density of the pyrolysis gas at the subcooling zone was calculated by following the same procedures discussed above. The result of the calculation is shown in Table H.5. The absolute viscosity and thermal conductivity of CO2 and CH4 were determined from gas property tables. The absolute viscosity and thermal conductivity of the gas mixture was calculated using Eq. (H.5) and (H.6), respectively. 2 2 4 4 G CO CO CH CH y y µ µ µ = + ( ) H.5 2 2 4 4 G CO CO CH CH k y k y k = + ( ) H.6 The values of the properties of the pyrolysis gas determined from the calculations above are summarized in Table H.5 and H.6. Table H.5: Pyrolysis Gas Properties Applied in Desuperheating Zone Property Symbol Numerical Value Unit Specific Heat at Constant Pressure cG,des 1,073.223 kJ/kg Density ρG,des 1.334 kJ/kg Absolute Viscosity μG,des 1.791 x 10 -5 kg/m·s Thermal Conductivity kG,des 2.501 x 10 -2 W/m·K Table H.6: Pyrolysis Gas Properties Applied in Subcooling Zone Property Symbol Numerical Value Unit Specific Heat at Constant Pressure cG,sub 1,027.952 J/kg·°C Density ρG,sub 1.505 kg/m 3 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 bo L sub V V = ( ) H.7 3 8 3 , 80 1 1 1 8.889 10 m s 15 min 60 1000 1000 L sub ml L m V s ml L ÷ | | | || | = = × | | | \ .\ .\ . 130 The flow area occupied by the bio-oil is L bo TP V A v = ( ) G.7 where: vTP = 0.125 m/s 8 7 2 8.889 10 7.111 10 m 0.125 bo A ÷ ÷ × = = × The flow area of the pyrolysis gas is 2 4 pg i bo i bo A A A d A t = ÷ = ÷ ( ) G.8 where: di = 0.0239 m ( ) 2 7 4 2 0.0239 7.111 10 4.479 10 m 4 pg A t ÷ ÷ = ÷ × = × 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 TP V A v = ( ) H.8 where: Apg = 4.479 x 10 -4 m 2 vTP = 0.125 m/s ( )( ) 4 5 3 , 4.479 10 0.125 5.599 10 m s G sub Q ÷ ÷ = × = × The void fraction in the subcooling zone was calculated using Eq. (H.9). , , , , G sub G sub L sub G sub V V V c = + ( ) H.9 5 , 8 5 5.599 10 0.998 8.899 10 5.599 10 G sub c ÷ ÷ ÷ × = = × + × 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 µ c µ µ c c µ | | | | \ . = | | ÷ + | | \ . ( ) H.10 1.505 0.998 976.709 0.492 1.505 1 0.998 0.998 976.709 sub x | | | \ . = = | | ÷ + | \ . 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 sub x x µ µ µ µ µ = + ÷ ( ) H.11 ( )( ) ( )( ) ( )( ) 3 , 1.505 976.709 3.050 kg m 0.492 976.709 1 0.492 1.505 TP sub µ = = + ÷ 131 The mass flux in the subcooling zone was calculated using Eq. (H.12). , , , , L sub G sub TP sub TP sub V V m A µ | | + = | | \ . - ( ) H.12 where: , L sub V = 8.889 x 10 -8 m 3 /s , G sub V = 5.599 x 10 -5 m 3 /s A = 4.486 x 10 -4 m 2 8 5 2 , 4 8.889 10 5.599 10 3.050 0.381kg m s 4.486 10 TP sub m ÷ ÷ ÷ | | × + × = = · | × \ . - As discussed in Section 3.7.2, the mass flux and quality are constant, that is, 2 , , , 0.381kg m s TP TP des TP con TP sub m m m m = = = = · - - - - 0.492 sub con des x 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 0.381kg m ×s 0.492 4.486 10 m TP m x A ÷ = = = × - The heat released by the volatiles in the desuperheating zone was calculated using Eq. (H.13). ( ) , 1 des TP G des des Q m A x h xc T ( = ÷ A + A ¸ ¸ - ( ) 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 ( )( ) ( )( ) ( )( )( ) 4 0.381 4.486 10 1 0.492 20,100 0.492 1, 073.223 10 des Q ÷ = × ÷ + ( ¸ ¸ 2.649 W des Q = The heat released by the bio-oil during condensation was calculated using Eq. (H.14). ( ) 1 con TP fg Q m x Ah = ÷ - ( ) H.14 where: hfg = 2,257,000 J/kg ( )( )( )( ) 4 0.381 1 0.492 4.486 10 2, 257, 000 con Q ÷ = ÷ × 195.949 W con Q = 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 c c xc x c = + ÷ ( ) H.15 ( )( ) ( )( ) ( )( ) , 1, 027.952 4,193.165 1, 666.313 J kg K 0.492 4,193.165 1 0.492 1, 027.952 TP sub c = = · + ÷ 132 , sub TP TP sub sub Q m Ac T = A - ( ) H.16 ( )( )( )( ) 4 0.381 4.486 10 1, 666.313 100 31 sub Q ÷ = × ÷ 19.668 W sub Q = 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 1 sub w w w w Q T T m c = + ( ) H.17 where w m 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 19.668 30 30.016 C 0.3 4,176 w T = + = ° 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 2 con w w w w Q T T m c = + ( ) H.18 ( )( ) 3 195.949 30.016 30.172 C 0.3 4,176 w T = + = ° Similarly, Tw4 was calculated using Eq. (H.19). 4 3 des w w w w Q T T m c = + ( ) H.19 ( )( ) 4 2.649 30.172 30.174 C 0.3 4,176 w T = + = ° 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 T T T T T T ÷ ÷ ÷ A = | | ÷ | | ÷ \ . ( ) H.20 ( ) ( ) , 110 30.174 100 30.172 74.715 C 110 30.174 ln 100 30.172 lm des T ÷ ÷ ÷ A = = ° ÷ | | | ÷ \ . ( ) ( ) 3 2 , 3 2 ln sat w sat w lm con sat w sat w T T T T T T T T T ÷ ÷ ÷ A = | | ÷ | | ÷ \ . ( ) H.21 ( ) ( ) , 100 30.172 100 30.016 69.906 C 100 30.172 ln 100 30.016 lm con T ÷ ÷ ÷ A = = ° ÷ | | | ÷ \ . 133 ( ) ( ) 2 , 1 , 2 , 1 ln sat w v ex w lm des sat w v ex w T T T T T T T T T ÷ ÷ ÷ A = | | ÷ | | ÷ \ . ( ) H.22 ( ) ( ) , 100 30.016 31 30 16.238 C 100 30.016 ln 31 30 ÷ ÷ ÷ A = = ° ÷ | | | ÷ \ . lm sub T 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 4 i o A D d t = ÷ ( ) H.23 ( ) ( ) 2 2 4 2 0.0351 0.0254 4.598 10 m 4 A t ÷ ( = ÷ = × ¸ ¸ 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 H w W m D Aµ = ( ) 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 0.023 Re 18, 727.60 4.598 10 8.030 10 w ÷ ÷ = = × × 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 0.4 Nu 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. Nu w w w H k h = D ( ) H.26 where: kw = 0.619 W/m·K ( )( ) 2 118.320 0.619 = = 3,179.614 W m K 0.023 w h · 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 des x x µ µ µ µ µ = + ÷ ( ) 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 10 1.464 10 kg m s 0.492 1.244 10 1 0.492 1.791 10 TP des µ ÷ ÷ ÷ ÷ ÷ × × = = × · × + ÷ × The Reynolds number at the desuperheating zone was calculated using Eq. (H.28). , , Re TP i TP des TP des m d µ = - ( ) H.28 where: TP m - = 0.381 kg/m 2 ·s di = 0.0239 m μTP,des = 1.464 x 10 -5 kg/m·s ( )( ) , 5 0.381 0.0239 Re 622.426 1.464 10 TP 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 k k 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 10 2.510 10 W m K 0.492 2.519 10 1 0.492 2.501 10 TP des k ÷ ÷ ÷ ÷ ÷ × × = = × · × + ÷ × 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 3.66 v,des i des TP,des h d = = k ( ) H.30 Rearranging Eq. (H.30) yields ( )( ) -2 2 3.66 2.501×10 = = 3.844 W m K 0.0239 v,des h · 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 i con v con m d µ = - ( ) H.31 where: TP m - = 0.381 kg/m 2 ·s di = 0.0239 m μv,con = 1.227 x 10 -5 kg/m·s 135 ( )( ) 5 0.381 0.0239 Re 742.777 1.227 10 con ÷ = = × Since Recon < 35,000, Eq. (H.32) was used to calculate the convection heat transfer coefficient during condensation. ( ) ( ) 1/4 1/4 3 sin 0.68 0.555 L,con L,con v,con L,con fg L,con sat i v,con L,con i sat i ρ ρ ρ g αk h + c T T h = 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 1/4 3 2 sin ln 0.68 ln 0.555 2 2 ln t i o w w,con L,con L,con vcon L,con v,in i i o i fg L,con sat i o i i t L,con i sat i t o w o i k T d h T + ρ ρ ρ g α k T T d d d h + c T T d d T = + k μ d T - T k d 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/m 3 hfg 2,257,000 J/kg ρv,con 0.590 kg/m 3 hw 3,179.614 W/m 2 · 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/s 2 cL,con 4,211 J/kg·K Applying bisection method to Eq. (H.33) yields, = 74.964 °C i T Substituting Ti back to Eq. (H.32) yields the convection heat transfer coefficient at the condensing zone, 2 , W m K v con h · =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 sub x x µ µ µ µ µ = + ÷ ( ) 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 10 3.196 10 kg m s 0.492 5.342 10 1 0.492 1.623 10 TP sub µ ÷ ÷ ÷ ÷ ÷ × × = = × · × + ÷ × The Reynolds number at the subcooling zone was calculated using Eq. (H.35). , , Re TP i TP sub TP sub m d µ = - ( ) H.35 where: TP m - = 0.381 kg/m 2 ·s 136 di = 0.0239 m μTP,sub = 3.196 x 10 -5 kg/m·s ( )( ) , 5 0.381 0.0239 Re 285.149 3.196 10 TP 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 k k 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 0.651 4.241 10 W m K 0.492 0.651 1 0.492 2.160 10 TP sub k ÷ ÷ ÷ × = = × · + ÷ × 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 3.66 v,sub i sub TP,sub h d = = k ( ) H.37 Rearranging Eq. (H.37) yields ( )( ) -2 2 3.66 4.241×10 = = 6.495 W m K 0.0239 v,sub h · 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. ( ) ln 1 1 Δ 2 o i lm i v t o w d d Q L= + + π T d h k d h ( ( ¸ ¸ ( ) H.38 For the desuperheating zone, ( ) ln 1 1 Δ 2 o i des des lm,des i v,des t o w d d Q L = + + π T d h k d h ( ( ¸ ¸ where: Qdes = 2.649 W ΔTlm,des = 74.715°C hv,des = 3.844 W/m 2 ·K ( ) ( )( ) ( ) ( ) ( )( ) ln 0.0254 0.0239 2.649 1 1 = + + 74.715 0.0239 3.844 2 204 0.0254 3,179.614 des L π ( ( ( ¸ ¸ = 0.123 m des L For the condensing zone, ( ) ln 1 1 Δ 2 o i con con lm,con i v,con t o w d d Q L = + + π T d h k d h ( ( ¸ ¸ where: Qcon = 195.949 W ΔTlm,con = 69.906°C 137 hv,con = 5,974.601 W/m 2 ·K ( ) ( )( ) ( ) ( ) ( )( ) ln 0.0254 0.0239 195.949 1 1 = + + 69.906 0.0239 5, 974.601 2 204 0.0254 3,179.614 con L π ( ( ( ¸ ¸ = 0.017 m con L For the subcooling zone, ( ) 1 1 Δ 2 o i sub sub lm,sub i v,sub t o w ln d d Q L = + + π T d h k d h ( ( ¸ ¸ where: Qcon = 19.668 W ΔTlm,con = 16.238°C hv,con = 6.495 W/m 2 ·K ( ) ( )( ) ( ) ( ) ( )( ) ln 0.0254 0.0239 19.668 1 1 = + + π 16.238 0.0239 6.495 2 204 0.0254 3,179.614 sub L ( ( ( ¸ ¸ = 2.488 m sub L The total condenser length is the sum of the lengths of each zone. des con sub L=L +L +L ( ) H.39 = 0.123+0.017+2.488 = 2.629 m L H.7. Pressure Drop Eq. (H.40) was used to calculate the pressure drop in each zone. ( ) 2 2 sin TP TP TP i TP f m L p g L d µ o µ A = ÷ - ( ) 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 Re TP TP f = ( ) H.41 The density of the two-phase mixture was calculated using Eq. (H.42). ( ) 1 G L TP L G x x µ µ µ µ µ = + ÷ ( ) H.42 For the desuperheating zone, , , 16 Re TP des TP des f = where: ReTP,des = 622.426 2 , 16 2.571 10 622.426 TP des f ÷ = = × The two-phase density is, ( ) , , , , , 1 G des L des TP des L des G des x x µ µ µ µ µ = + ÷ 138 where the values of ρG,des and ρL,des are shown in Table H.1 and H.5, respectively. ( )( ) ( )( ) ( )( ) 3 , 1.334 0.5816 0.805 kg m 0.492 0.5816 1 0.492 1.334 TP des µ = = + ÷ The pressure drop at the desuperheating zone is ( ) 2 , , , 2 sin TP des TP des des TP des des i TP des f m L p g L d µ o µ A = ÷ - ( )( ) ( ) ( )( ) ( )( )( ) ( ) 2 2 2 2.571 10 0.381 0.123 9.81 0.805 0.123 sin 20 0.0239 0.805 des p ÷ × A = ÷ ÷ ° 0.380 Pa des p A = The pressure at the desuperheating zone exit is 2 1 186.259 0.380 185.879 Pa (gage) des p p p = ÷A = ÷ = 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 Re TP con TP con f = where: ReTP,con = 742.777 2 , 16 2.154 10 742.777 TP con f ÷ = = × 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 µ µ µ + = where the values of ρv,con and ρL,con are shown in Table A10.7. 3 , 0.590 958.320 479.455 kg m 2 TP con µ + = = The pressure drop at the condensing zone is, ( ) 2 , , , 2 sin TP con TP con con TP con con i TP con f m L p g L d µ o µ A = ÷ - ( )( ) ( ) ( )( ) ( )( )( ) ( ) 2 2 2 2 2 5.611 10 0.381 1.743 10 9.81 479.455 1.743 10 sin 20 0.0239 479.455 con p ÷ ÷ ÷ × × A = ÷ × ÷ ° 28.038 Pa con p A = The pressure at the condensing zone exit is 3 2 185.879 28.038 157.841Pa (gage) con p p p = ÷A = ÷ = 139 For the subcooling zone, , , 16 Re TP sub TP sub f = where: ReTP,sub = 285.149 2 , 16 5.611 10 284.149 TP sub f ÷ = = × The two-phase density is, ( ) , , , , , 1 G sub L sub TP sub L sub G sub x x µ µ µ µ µ = + ÷ where the values of ρG,sub and ρL,sub are shown in Table H.3 and H.6, respectively. ( )( ) ( )( ) ( )( ) 3 , 1.505 976.709 3.050 kg m 0.492 976.709 1 0.492 1.505 TP sub µ = = + ÷ The pressure drop at the subcooling zone is, ( ) 2 , , , 2 sin TP sub TP sub sub TP sub sub i TP sub f m L p g L d µ o µ A = ÷ - ( )( ) ( ) ( )( ) ( )( )( ) ( ) 2 2 2 5.611 10 0.381 2.488 9.81 3.050 2.488 sin 20 0.0239 3.050 sub p ÷ × A = ÷ ÷ ° 26.027 Pa sub p A = The pressure at the subcooling zone exit is 4 3 157.841 26.027 131.814 Pa (gage) sub p p p = ÷A = ÷ = 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. 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