Optimizing a Swirl Burner for Pyrolysis Liquid Biofuel (Bio-Oil

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Optimizing a Swirl Burner for Pyrolysis Liquid Biofuel (Bio-Oil) Combustion without Blending Sina Zadmajid, Steven Albert-Green, Yashar Afarin, and Murray J. Thomson Energy Fuels, Just Accepted Manuscript • DOI: 10.1021/acs.energyfuels.6b03417 • Publication Date (Web): 13 Apr 2017 Downloaded from http://pubs.acs.org on April 14, 2017

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Optimizing a Swirl Burner for Pyrolysis Liquid Biofuel (Bio-Oil) Combustion without Blending Sina Zadmajid, Steven Albert-Green, Yashar Afarin, and Murray J. Thomson* Department of Mechanical and Industrial Engineering, University of Toronto, 5 King’s College Road, Toronto, Ontario, Canada M5S 3G8

ABSTRACT: Bio-oil, also called pyrolysis liquid biofuel, is made from the pyrolysis of waste biomass and provides a carbon-neutral alternative fuel. This study focuses on optimizing swirl burners for use with bio-oil. Spray combustion of bio-oil/ethanol blends was previously studied using a 10 kW swirl burner. The previous burner had a small combustion chamber with no refractory lining. It was not feasible to stabilize flames of bio-oil without blending with ethanol and the pollutant emissions were relatively high. In this study, the burner is upgraded by implementing a refractorylined combustion chamber and increasing the size of the chamber to investigate the relationship between the burner design and combustion performance of bio-oil. The main reasons for upgrading the burner are to eliminate ethanol addition to bio-oil and stabilize the pure bio-oil flame, achieve lower pollutant emissions, and make the burner more comparable to industrial bio-oil burners. In addition to flame stability, fuel boiling inside the nozzle and nozzle coking are two other challenges that require some optimization and adjustments in the setup. After modifying the new burner configuration and adjusting the operating parameters, the pure bio-oil flame is stabilized with minimal nozzle coking, and gaseous exhaust emissions. The new burner has a significantly lower amount of heat loss and hence, a higher gas temperature within the combustion chamber as compared to the previous one. The effects of operating conditions on the gaseous pollutants are also investigated. CO and unburned hydrocarbon emissions are extremely sensitive to the operating parameters such as equivalence ratio, swirl number, and atomizing air flow rate. Moreover, the pilot flame energy and primary air preheat temperature play important roles in ignition quality of bio-oil, flame stability, and nozzle coking. The dominant mechanism for NOx emissions is the conversion of fuel-bound nitrogen. The new burner has also decreased particulate matter emissions as compared to the previous burner by achieving higher temperatures, which favors burnout of char particles within the flame.

1. INTRODUCTION Wood-derived pyrolysis liquid biofuel, also called bio-oil, is produced from fast pyrolysis of waste biomass. Fast pyrolysis process is an inexpensive and environmentally attractive method of converting biomass into a liquid fuel, because it can be a self-sustaining process.1,2 Bio-oil is carbon neutral and has the potential to replace petroleum fuels in some applications such as industrial-scale boilers and furnaces,3-5 home-scale heating applications,6 gas turbines,7,8 and compression ignition diesel engines.9,10 In contrast to many biofuels such as ethanol, which are made from food resources, bio-oil is made from waste biomasses and therefore has no impact on food supply, making it a sustainable source of energy.11-13 ASTM has created a standard specification for pyrolysis liquid biofuel (ASTM D7544-09). It contains significant water and oxygen, which are jointly responsible for its poor ignition quality, comparatively lower energy density than hydrocarbon fuels, and reduced flame stability.14,15 Several studies have shown that bio-oil has a multiphase structure. It consists of an aqueous phase, which contains most of the low molecular weight (LMW) compounds, as well as a non-aqueous phase, which is characterized by high molecular weight (HMW) compounds and oligomers.16 Bio-oil also contains solid content, including organic char particles and inorganic ash, which can clog filters and the nozzle while forming particulate matter emissions. This biofuel has a wide volatility range, high surface tension, and high viscosity, which are important with respect to spray atomization. The viscosity of bio-oil strongly depends on its temperature, which is critical for fuel pumping and 1

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injection. In order to decrease its viscosity and improve the atomization quality, bio-oil is usually preheated up to about 80°C prior to injection.10,17 Bio-oil droplet combustion is characterized by two major stages.17 First, most of the volatile compounds burn in a homogeneous combustion mode. Fuel remnants following the first stage are known as secondary char particles, which is composed of the primary char found in bio-oil and some polymerized high molecular weight (HMW) material formed during the combustion first stage. In the second stage of combustion, the char particle burns slowly in a heterogeneous mode, resulting in particulate matter (PM) emissions. With sufficient residence time, PM should be almost entirely composed of ash. The replacement of petroleum oils with bio-oil in industrial-scale applications is a relatively novel technology. While bio-oil is a promising candidate to replace conventional fuels in many stationary applications such as boilers and furnaces, their physical properties and chemical composition differ. In order to increase pyrolysis liquid usability, three main techniques can be used: upgrading bio-oil fuel quality,18-20 co-combustion of bio-oil with fossil fuels,15,21 and optimizing combustion devices.10,22,23 The combustion of bio-oil/ethanol mixtures at different blend ratios has been previously studied in our research group with a 10 kW swirl burner.17,24-27 In this study, modifications are applied to this burner in order to make it a more representative of actual industrial burners and also to have a stable flame using pure bio-oil. Two refractory linings are installed to reduce the amount of heat loss from walls and also to have more uniform temperature distribution inside the combustion chamber. Furthermore, to prevent the unburnt fuels from impinging on the walls, the diffuser is removed and the volume of the combustion chamber is increased in the new design. These modifications decrease the amount of heat loss, improve ignition quality, and enhance flame stability. The new setup addresses common problems associated with using bio-oil in small-scale burners, the most critical of which include: ignition characteristics, flame stability, nozzle coking, and pollutant emissions. 2. EXPERIMENTAL METHODOLOGY 2.1. Burner Assembly. 2.1.1. Previous Burner Assembly. A 10 kW swirl burner, designed by Tzanetakis et al.,24 was previously used to study bio-oil combustion. The burner was constructed from 316-grade stainless steel to counter the corrosive nature of the fuel. A variable swirl generator was installed on top of the combustion chamber, and the fuel was injected downwards by an internally-mixed air-blast atomizing nozzle into the combustion chamber. There was no refractory lining inside the combustion chamber and hence, the amount of heat loss was significant. 2.1.2. Upgraded Combustion Chamber. In this study, a new refractory lined combustion chamber was designed and manufactured to replace the previous one. It has a larger combustion chamber with two refractory linings in order to decrease heat loss and improve combustion efficiency. The upgraded assembly utilizes the existing swirl box as the swirl air generator. A door is available for maintenance and cleaning of the combustion chamber, fuel nozzle, and viewports after each test. Figure 1 depicts a schematic of the new burner assembly and its overall dimensions. Similar to industry-scale burner designs, the upgraded combustion chamber does not have a diffuser section.

Figure 1. Schematics of the new burner assembly and the refractory linings 2


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As shown in Figure 1, there are two refractory linings used in the new combustor (3” thick Alumina-Silica fibers and 1.5” thick Durablanket S insulation), which improve the energy efficiency of the burner and significantly decrease the amount of heat loss through the burner walls. The design target was to decrease the outer wall temperature of the new burner to below 60°C as compared to 500°C of the previous one. This has been achieved as the outer wall temperature is below 50°C during the experiments. 2.1.3. Variable Swirl Generator. Turbulence is an important factor in non-premixed flames, because it increases mixing between air and fuel, enhancing ignition and flame stability. In order to promote turbulence, a movable block type swirl generator is used to create a swirling flow around the fuel nozzle.24 A swirl flow incorporates a certain amount of angular momentum and velocity, which can induce radial and axial pressure gradients, with low pressure zones in the center and upstream of the flow. This creates a central recirculation zone (CRZ) inside the combustor which heats incoming fuel, assisting ignition and lengthens burner residence time in order to burn out particulates. The swirl box used in this study consists of 8 movable swirl blocks, which can generate a swirl number from 0 to 5.41. The details regarding swirl number calculations and analysis for movable block geometries can be found elsewhere.24,28,29 However, the predicted swirl number does not remain constant between the swirl block exit plane and the nozzle tip. Others have proved that the swirl number can decay up to 40% along the annular region in the burner throat.28 Hence, these swirl numbers are only used as relative indicators of the swirl intensity inside the burner. 2.1.4. Fuel Atomizing Nozzle. The nozzle assembly includes a liquid cap, an air cap, and a ¼” JPL back-connect body provided by BEX Engineering Ltd, as well as a 13-inch-long extension tube. All these parts are constructed from 316grade stainless steel due to the corrosive nature of the fuel.30 Four different nozzles are studied in the experiments, all of which are internally-mixed air-blast atomizing nozzles, which differ only with respect to the liquid and air cap designs. The first nozzle used in the experiments is a JPL26B nozzle assembly from BEX Engineering Ltd, including a JL40100 liquid cap and a JPG60 air cap. As shown in Figure 2, the liquid cap has a discharge orifice of 1 mm and the air cap has six discharge orifices of 0.94 mm each, creating a hollow cone spray pattern. This nozzle is the same as the one used in the previous burner. However, in order to address the nozzle coking problem (section 3.5), three other nozzles are utilized to find the best design for bio-oil combustion.

Figure 2. Schematic of the atomizing nozzle no.131 The second nozzle is a JPL16 nozzle assembly, also from BEX Engineering Ltd, including a JL2050 liquid cap and a JPG15 air cap. This nozzle is smaller than the first one, with an orifice diameter of 0.51 mm on both the liquid and air caps. The next two nozzle assemblies use the same liquid cap as the first nozzle (JL40100 liquid cap), but utilize different JPG60 air cap orifice designs. One of the air caps has an orifice diameter of 0.80 mm, whereas the other one has an orifice diameter of 0.66 mm. Table 1 summarizes the nozzle specifications. Table 1. Nozzle specifications Liquid cap model no.

dL (mm)

Air cap model no.

dA (mm)

1

JL40100

1

JPG60

0.94

2

JL2050

0.51

JPG15

0.51

3

JL40100

1

JPG60 custom ma-

0.80

Nozzle no.

3


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chined 4

JL40100

1

JPG60 custom machined

0.66

Bio-oil is difficult to ignite and has poor flame stability and so it is important to use a proper nozzle design. Different nozzle designs produce different droplet size ranges, which can significantly affect the ignition quality of the bio-oil. Literature suggests that the Sauter mean diameter (SMD) of an air-blast atomizer can be estimated by this correlation:32 .

SMD 0.48

.

. 1 0.15

1

(1)

where is the liquid orifice diameter, is surface tension, and are densities of air and liquid, is the relative velocity of air and liquid, and are mass flow rates of air and liquid, and is the dynamic viscosity of liquid at the nozzle temperature. It should also be noted that this droplet size correlation is not designed for an internally mixed atomizer and it is therefore only being used for estimations of atomization trends. When some components of the liquid fuel reach their boiling point (80-90°C) and evaporate inside the nozzle, some bubbles start to form within the liquid stream and expand as the liquid undergoes a pressure drop, causing the surrounding liquid to shatter and thus atomize violently, and creating random instabilities. This refers to a flashing phenomenon,33 which can cause flame instabilities and blowout. To avoid fuel from boiling inside the nozzle, its temperature (at the location shown in Figure 3) is maintained around 65-75°C prior to injection with a cooling system designed as shown in Figure 3, which is similar to that of the previous burner. A stainless steel tube with outer and inner diameters of 1/16” and 0.04”, respectively, is wrapped around the nozzle tube. A needle valve is used to control the flow rate of cold water passing through the cooling tube, with a maximum flow rate of 72 ml/min.

Figure 3. Nozzle cooling system 2.1.5. Ignition System. A continuous methane/oxygen pilot flame is used to ignite and stabilize the bio-oil flame within the burner. It provides only 3% of the total energy input (0.3 kW) at flow rates of 0.548 L/min methane and 1.121 L/min oxygen. The pilot consists of a hoke jeweler's soldering torch body (Hoke model No. 110-406, provided by Contenti Co.), a 20” long extension stainless steel tube (1/4” OD), and a pilot tip. Originally, the pilot tip was a standard No. 7 brass tip with a 1.2 mm orifice diameter. However, due to high temperatures achieved within the new burner, the brass tip melted. Thus, a Hi-Heat torch tip (Item No. 14.157, provided by Grobet USA) is used for the pilot flame to run the premixed CH4/O2 flame in order to avoid melting. With the new design, an electric igniter is used to start the pilot flame directly from within the burner to avoid opening the door or removing the pilot from its port. The pilot flame is ignited by moving its tip close to the electric igniter tip inside the burner and pushing a red button to generate a spark. Figure 4 demonstrates the ignition system. This setup is safer and allows the operator to reignite the pilot flame in case of extinction during the test.

4


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Figure 4. Schematic of the ignition system 2.2. Fuel Analysis. A basic fuel analysis (i.e. water, solid, and ash contents) is provided from the manufacturer, whereas other important properties of bio-oil, namely, the elemental composition, kinematic viscosity, and the gross heat of combustion are measured before conducting the experiments. An elemental analysis determines carbon, hydrogen, and nitrogen contents of fuel, which is done in the ANALEST facility at the University of Toronto according to standard ASTM D5291. The oxygen content of fuel is then calculated by the difference. The kinematic viscosity of fuel is measured in the Rheology Laboratory at the University of Toronto, which is done according to standard ASTM D445 using a viscometer placed in a constant temperature tank. The gross heat of combustion is measured in the Pulp and Paper Center at the University of Toronto, using a Parr 6300 calorimeter according to standard ASTM D240. In this method, the sample is weighed and placed inside the calorimeter. Before taking measurements, a sample of pure ethanol, which has a known heating value (HHV = 29.685 MJ/kg), was used as a reference to confirm the accuracy of the results. The average error was found to be ±2.4%. 2.3. Overall Experimental Setup. The experimental setup includes various inputs, outputs, and sample lines, which are depicted in Figure 5.

Figure 5. Schematic of the experimental setup Two peristaltic pumps are used to provide an adequate fuel flow rate and inject bio-oil into the combustor by the atomizing nozzle. The required fuel flow rate to reach the 10kW energy throughput is calculated based on the lower heating value (LHV) of the fuel. This energy input remains constant for all burner tests to ensure that they are comparable. The fuel flow rate for pure bio-oil is about 30.5 mL/min in this study. Although this energy throughput is lower than industrial furnaces, it makes it possible to study different operating conditions for bio-oil combustion and their trends without using a significant amount of fuel. Moreover, this work can be directly applied to smaller residential boilers or any other small-scale burner application that utilizes pyrolysis liquids. A stack fan is used at the end of the exhaust line to pull room air into the combustor and provide the primary air for combustion putting the exhaust line and combustion chamber both under a slight vacuum. An air heater is used to preheat the primary air and subsequently, to preheat the fuel inside the nozzle. The pilot flame is used to ignite the fuel and stabilize the main flame. The exhaust products are collected, and the gas phase emissions as well as the particulate matter are measured to investigate the relationship between the burner design parameters and combustion performance of the bio-oil. 5


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2.4. Flame Visualization. Two 6.5” x 6.5” x 1/4” thick clear fused quartz plates are used in the combustor's two viewports in order to provide direct visualization of the flame, which was not possible in the previous burner. A 10 megapixel Kodak (EasyShare Z1012 IS) digital camera is mounted on a tripod and placed in front of one of the quartz viewports to capture pictures and videos from the flame. Figure 6 shows a picture taken from the viewport, identifying the major visible constituents of the combustor and bio-oil flame.

Figure 6. Image of the flame and combustor through the viewport 2.5. Gas Phase Emission Measurement. Using heated sample lines, a heated filter, and three measurement instruments, gas phase pollutants as well as the percent oxygen in the exhaust products are detected. A flame ionization detector (FID) is used to measure the amount of unburned hydrocarbons (UHCs) in the exhaust sample. The UHC emissions are reported as ppm of methane with an uncertainty of ±3 ppm. A Nicolet 380 Fourier transform infrared spectrometer (FTIR) is used to measure gaseous pollutants such as CO, NOx, methane (CH4), formaldehyde (CH2O), and acetaldehyde (C2H4O). The percent of oxygen presented in the exhaust is continuously measured by a Zirconia (ZrO2) model OXY6200 oxygen sensor. The equivalence ratio is then back-calculated assuming complete combustion takes place and the measured O2 in the exhaust is used. Additional details regarding the principle of operation of these instruments can be found in previous works.24,34 2.6. Particulate Matter Measurement and Analysis. 2.6.1. Isokinetic Particulate Sampling. In order to obtain a representative sample of exhaust PM, isokinetic conditions are required. This is achieved when a stream of gas enters a sampling probe parallel to the flow stream with no change in velocity.35 In this study, a null-type isokinetic sampler (∆P=0) is used.36 Additional details about the PM sampling system used in this study can be found elsewhere.26 Using a 47 mm diameter Tissuquartz filter (Product No. 7202, provided by Pall Life Sciences), PM is collected for analysis. A vacuum pump is located at the end of the sampling line to pull the exhaust gases through the filter. The sampling time is usually less than 5 minutes to avoid overloading the filter, which can then clog the line. In order to replace the filter with a new one, a valve closes the main sampling line while another valve is opened to put the flow through a bypass line. A condenser (Part No. 2151414, provided by Seakamp Engineering Inc.) is placed before the vacuum pump to cool down the exhaust gases and remove water from them prior to reaching the pump. The flow rate of the gas exiting the condenser (dry gas) as well as its temperature and pressure are measured in order to calculate the actual sampling flow rate (!"#$%&'( ): !"#$%&'( +"#, ⁄+,-,#& / 1 (2) 0123 / 1 !$'#")*'(

Here, !$'#")*'( is the flow rate of dry gas measured by a rotameter prior to the vacuum pump, +"#, is the saturation pressure at the temperature of dry gas, +,-,#& is the absolute pressure in the line measured by the gauge in the condenser, and 012 3 is the molar fraction of water in the exhaust gases. 2.6.2. Gravimetric Analysis and Loss on Ignition. While the burner is running at a steady operating condition, two consecutive filters are used to collect PM samples. These filters are weighed and stored in petri dishes before and after each test to find the total amount of PM collected, using the following equation: PM

∆67&,'* !"#$%&'( 6)'& ∆8 !,-,#& 6


(3)

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Here, ∆67&,'* is mass difference between the blank and loaded filter, 6)'& is mass flow-rate of fuel, and !"#$%&'( ⁄!,-,#& is the ratio of the sampling flow-rate to the total exhaust flow-rate. !"#$%&'( is found from equation (2) and !,-,#& is calculated assuming complete combustion and using the equivalence ratio obtained from the oxygen sensor. PM consists of partially-burned carbonaceous residues (CR), which result from incomplete combustion, and ash. In order to measure the amount of each fraction, a loss on ignition test, which is an ASTM D4422-03 standard method, is performed after collecting the PM. Prior to a PM sampling test, blank filters are first placed inside a muffle furnace at 750°C in order to burn out any possible particles accumulated on them. The filters are then used in the burner to collect PM. After finishing the test, the filters are dried at a temperature of 150°C in the muffle furnace for two hours. At this stage, the mass difference indicates the amount of water collected on the filter, and the mass difference between the dried filter and its initial condition indicates the total PM collected. Afterwards, the temperature of the muffle furnace is increased to 750°C for one hour to burn off any possible carbonaceous residue on the filter. The mass difference after this stage represents the amount of CR and from knowing the mass of total PM and CR, the amount of ash collected on the filter can then be calculated. 2.6.3. Uncertainty Analysis. According to equation (3), PM is a function of several variables, each of which is associated with a particular uncertainty. Therefore, the uncertainty of the results (9:; ) can be measured by the following differential equation: 9:;

B >+? >+? >+? >+? ±[ 9∆$ 9$ 9∆, 9* ] >∆ > >∆8 >@

(4)

where 9∆$ , 9$ , 9∆, , and 9* are the uncertainties associated with the filter weight difference, mass flow rate of fuel, sampling time, and the ratio of !"#$%&'( /!,-,#& , respectively. 2.7. Heat Exchanger Analysis. The total energy input of the burner is derived from the energy of the fuel, the pilot flame, and the air heater. The output energy is the sum of the heat extracted from burner (D'E,* ) and the amount of heat loss. Following equations show the energy balance of the system: D7F D-),

D6)'& D%7&-, DG'#,'* D'E,* D&-""

(5) (6)

The heat extracted from the burner includes the heat released from the exhaust gases (DH#" ) and also the heat released from the condensation of water vapor (DI-F( ) within the exhaust gases. This energy is extracted by a watercooled heat exchanger (supplied by Polar Power Inc.), which uses cold tap water at a temperature of around 7 to 14°C to extract the energy and cool down burner the exhaust gases. The energy extracted by the heat exchanger results in a change in enthalpy of the cooling water. Therefore, the flow rate of the cooling water as well as the temperature difference of water entering and exiting the unit are measured in order to calculate the heat extracted from the burner, which is shown in the following equation: D'E,* DH#" DI-F( J . ΔℎJ J M% ΔNJ

(7)

Here, J is the mass flow rate of cooling water, ΔℎJ is the change in enthalpy of water, and M% is the constantpressure specific heat capacity of water. 2.8. Burner Test Procedure. Around 1 to 2 hours before each test, all the sample lines are warmed up, the FID, pilot flame, atomizing air and cooling water for the heat exchanger are turned on and room air is drawn into the burner by the stack fan. Initially, the burner is warmed up with ethanol (EtOH) for about 30 minutes before the fuel line is switched to bio-oil at the desired flow rate, which is calculated based on the heating value of the fuel to reach 10 kW. After reaching a steady state condition, all the measurements are taken. Steady state condition is defined as the time when the exhaust temperature gradient is close to zero (below 2°C/min). At the end of each test, the burner is flushed with EtOH to clean the fuel lines. 7


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3. RESULTS AND DISCUSSION 3.1. Fuel Analysis. Table 2 summarizes the properties of the three bio-oil batches used in this study with their typical ranges suggested in the literature. The bio-oils used in this study were all made of waste white-wood rather than bark and branches, and for that reason, they contained very low amounts of fuel bound nitrogen that were below detection thresholds of the elemental analysis (below 0.01 wt %). Table 2. Properties of the bio-oil used throughout this study

Unit

Bio-oil batch 1a

Bio-oil batch 2b

C-H-O-N

wt %

43.686.7549.570.00

Water

wt %

Solids

Parameter

Bio-oil

Typical bio-oil

batch 3c

(from literature)d

45.807.3846.820.00

45.627.3047.08-0.00

55-6-38-0.15

23.36

22.7

21.8

15 to 30

wt %

0.21

0.14

0.07

0.2 to 1

Ash

wt %

0.12

0.13

0.09

0 to 0.3

Density (20°C)

kg/m3

1180

1125

1171

1100 to 1300

Gross Heat of Combustion (HHV)

MJ/kg

17.99

18.31

18.43

16 to 23

LHV

MJ/kg

16.56

16.74

16.87

15.6 to 21.6

cSt

42.2

N/A

35.4

10 to 100

Kinematic Viscosity (40°C) a

Batch 1 was used to address initial flame instabilities within the new burner and find the best burner configuration. b Batch 2 was used to address flash-atomization instabilities, nozzle fuel coking, the sensitivity analysis of primary air flow rate and swirl number, and the heat exchanger analysis. c Batch 3 was used to study gas phase emissions during transient base point operation to complete the parametric sensitivity analysis (for atomizing air flow rate, pilot flame energy, primary air and fuel preheat temperature), and to study particulate matter emissions. d Taken from various literature sources.15,19,22,37,38 The gross heat of combustion or higher heating value (HHV) was directly measured for batches 2 and 3; however, it was not possible to directly measure the HHV of batch 1 using the calorimeter, because it could not ignite. Hence, the HHV of different mixtures of ethanol and bio-oil were measured and the HHV of the batch 1 bio-oil was determined by extrapolating down to zero ethanol. The HHV of bio-oil can also be calculated from equation (8) as suggested by Demirbas et al.39 for wood-derived biofuels such as bio-oil. HHV 33.5 × [C] 142.3 × [H] / 15.4 × [O] / 14.5 × [N]

;W XH

(8)

Here, [C], [H], [O], and [N] are concentrations of carbon, hydrogen, oxygen and nitrogen in the fuel. This correlation was previously used to calculate HHVs for 11 similar bio-oil batches from the same producer.40 The comparison between the calculated HHVs and the measured values for the three bio-oil batches used in the present study indicated an average error of 2.3%. It should be noted that the standard deviation of these errors was 2.5%. Thus, the equation has the same accuracy as direct measurements with the calorimeter. Figure 7 plots the viscosity of batch 3 at different temperatures, showing the important inverse relationship between temperature and viscosity. Preheating bio-oil reduces its viscosity and therefore, improves atomization and subsequently its ignition quality. Therefore, an air heater is used to preheat the primary air and ultimately the bio-oil inside the fuel nozzle prior to injection. 8


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Energy & Fuels 100 80 Viscosity (cSt)

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

60 40 20 0 20

30

40

50 60 70 Temperature (ºC)

80

90

Figure 7. Viscosity of bio-oil batch 3 versus temperature 3.2. Flame Stability. Flame instabilities create significant smoke emissions and particulate formation as well as substantial fluctuations in the air flow within the burner due to pressure changes. Ethanol addition to bio-oil increases the heating value of the fuel and its volatility while also greatly decreasing its viscosity, improving fuel atomization, ignition quality, and ultimately flame stability. Therefore, bio-oil/EtOH blends were used during initial experiments to address flame instabilities and modify the burner configuration in order to get stable flames. Despite achieving stable flames of 50/50 bio-oil/EtOH (by vol.) and 60/40 bio-oil/EtOH in the new burner configuration (Figure 8), the flame of 80/20 bio-oil/EtOH was unstable and was not anchored to the nozzle, as shown in Figure 9.a. A low degree of primary air swirl was the likely source of flame instabilities.

Figure 8. Schematic of the new burner assembly with the swirl outlet pipe Several approaches were used to address these instabilities and to find the best burner configuration. One method for improving flame stability was to remove the “swirl outlet pipe” (shown in Figure 8) to decrease the axial distance between the swirl blocks and the fuel nozzle. As mentioned earlier in section 2.1.3, swirl decays along the annular region in the burner throat so this modification mitigated that effect. An image of the flame after removing this pipe is shown in Figure 9.b. Although the flame became more stable and more anchored to the nozzle than it was with swirl outlet pipe, it was asymmetrical and the ignition quality was poor. The next solution was to place a duct reducing sleeve inside the burner throat, as shown in Figure 10, in order to reduce the cross-sectional area of the throat. This metal sleeve also created a circular area inside the burner throat which previously had an irregular shape due to the insulation linings inside the throat. As the result, the air flow around the nozzle and hence, the flame jets became more symmetrical than before (Figure 9.c). Therefore, this configuration was used to burn pure bio-oil for the remainder of this study.

9


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(a) burner assembly with swirl outlet

(b) After removing the swirl outlet

pipe

pipe

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(c) After placing sleeve inside burner throat

Figure 9. Flame stability: flames of 80/20 bio-oil/EtOH at various burner configurations, with flame (c) representing the desired stable flame.

Figure 10. A cross-sectional view of the final burner configuration with its burner throat with the added sleeve 3.3. Burning Pure Bio-Oil. With the previous burner, it was necessary to mix bio-oil with at least 20% ethanol (by volume) in order to obtain stable flames. Although the previous burner was capable of operating with a mixture of 90/10 bio-oil/ethanol or pure bio-oil, the flame stability was very poor and would easily extinguish, resulting in high UHC emissions exceeding 150 ppm.24 However, the new combustion chamber (Figure 10) has made it possible to burn pure bio-oil because of the comparably stronger recirculation zones around the nozzle and the higher temperatures within the combustion chamber as compared to the previous one. Different flames of bio-oil/EtOH blends are compared in Figure 11. As explained earlier in section 1, char burning is indicative of bio-oil flames, which is most evident in the pure bio-oil flame (Figure 11.d).

(a) Pure ethanol

(b) 80/20 biooil/EtOH

(c) 90/10 biooil/EtOH

(d) Pure bio-oil

Figure 11. Flame images of pure ethanol, 80/20 bio-oil/EtOH, 90/10 bio-oil/EtOH, and pure bio-oil 3.4. Fuel Boiling within the Nozzle. After flames of pure bio-oil stabilize, the burner temperature continually increases until reaching a steady state condition. Thermal feedback from the flame to the nozzle causes the fuel temperature to also increase until after only a few minutes, the fuel starts to boil inside the nozzle cap, causing flashing. Although the increased fuel temperature reduces viscosity, fuel boiling causes severe instabilities and flame blow out, as explained in section 2.1.4. Figure 12 illustrates this phenomenon by showing images of a flashing bio-oil flame with a time interval of 0.1 s between images. Hence, it is important to control the fuel temperature inside the nozzle and cool down the fuel line when its temperature exceeds 70-80 °C. The cooling system, which is described in section 2.1.4, is 10


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Energy & Fuels

the best means to cool down the fuel within the fuel line. Although the cooling system decreased the fuel temperature in the nozzle extension tube, it was found that this cooling system was unable to prevent fuel from boiling inside the nozzle cap.

a

b

c

d

e

f

g

h

i

j

k

l

Figure 12. Flashing instabilities: from (a) to (l), images of the flame with a time interval of 0.1s between images. The fuel boiling problem was first addressed by changing many operating parameters such as the atomizing air flow rate, swirl number, primary air preheat temperature, pilot flame energy and positioning, primary air flow rate, and fuel flow rate. However, these techniques were found to be ineffective, and another method had to be employed to keep the fuel within the nozzle from over-heating. Therefore, a 1/4" thick flexible ceramic insulation sheet was wrapped around the nozzle cap to insulate it against the intense heat, as shown in Figure 13. This sheet can withstand temperatures up to 1100°C. With this technique, the fuel boiling problem was solved and the nozzle cooling system was then used effectively in combination to control the fuel temperature.

Figure 13. Nozzle insulation (the nozzle was moved further down the burner throat to take this picture)

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3.5. Nozzle Fuel Coking and Clogging. Bio-oil can polymerize either inside the nozzle and clog the fuel line or on the external surface of the nozzle's air cap and form large carbonaceous residues, also known as coke.4,6,41 It was found that during each test and almost 30 minutes after initiating the bio-oil flame, large amounts of coke would start to build up on the external surface of the nozzle cap (nozzle no.1). The bio-oil sprayed out of the nozzle polymerized immediately on the hot external surface of the nozzle cap before evaporating and burning out. Figure 14 depicts this nozzle cap after a typical burner test.

Figure 14. Carbonaceous material (coke) formed on the external surface of nozzle no.1, image follows a combustion test Coking results in poor atomization and poor combustion indicated by high UHC and CO emissions. Similar to the previous section, several techniques were employed to decrease the degree of coking on the nozzle surface. The high temperature of the surface of the nozzle's air cap was one of the major reasons for the observed coke formation, mainly because of the hot gas recirculation zones around the nozzle. Decreasing the swirl number was one way to control the coking tendency of nozzle; however, decreasing the degree of swirl would also reduce flame stability. Despite this, the swirl number was reduced from the maximum value (S = 5.41) to S = 3.38 which allowed for a stable flame with a reduced coking tendency; though after an hour of operation, major coke formation on the nozzle still occurred. An effective solution was to decrease the droplet size since smaller droplets require a shorter residence time for burnout.42 According to equation (1), the droplet size of an atomized liquid depends on various parameters, particularly the atomizing air flow rate. However, there is a limited range for the atomizing air flow rate when using pure bio-oil, due to the tendency of the flame to blow out. In general, the atomizing air flow rate range for flame stability is narrower for a less volatile fuel such as bio-oil.42 For nozzle no.1, the atomizing air flow rate range reaches a minimum at 7 SLPM, below which bio-oil is not atomized properly and starts to dribble out of the nozzle, and a maximum flow rate at 15.5 SLPM, above which the flame jets become detached and the flame eventually blow out. Considering this limitation, increasing the atomizing air flow rate was not feasible to solve this problem. The orifice size of the nozzle is another parameter that affects the droplet size and SMD value. Therefore, in order to decrease the droplet size, three smaller nozzles were tested to address this coking problem. The schematic of air-blast atomizing nozzles and the dimensions of their discharge orifices can be found in Figure 2 and Table 1, respectively. Nozzle no.2, which uses the smallest liquid and air caps, was the first replacement nozzle tested. Using this nozzle, the fuel was ejected further into the combustor and far from the air cap surface due to the higher pressure within the nozzle; therefore, there was no coke formation on the external surface of the air cap. Although this nozzle solved the external coking problem, the small (0.51mm) liquid outlet within the nozzle clogged after 40 minutes of running on bio-oil and the test had to be terminated. Small orifices are far more prone to blockage from solid build-up or fuel polymerization. The discharge orifice of nozzle no.1, on the other hand, is large enough to avoid clogging. With these findings in mind, nozzles no.3 and no.4 were custom machined. These nozzles use the same liquid cap (JL40100) and a similar air cap (JPG60) as nozzle no.1, but with different air cap discharge orifice diameters. Nozzle no.3 has smaller discharge orifices than nozzle no.1, but it has larger discharge orifices than nozzle no.2. As it was expected, nozzle no.3 was found to have a less coking tendency than nozzle no.1. However, a further decrease in the discharge orifice size of the nozzle slightly increased the amount of coke such that nozzle no.4, which uses smaller discharge orifices than nozzle no.3, showed a slightly higher coking tendency than nozzle no.3. Therefore, nozzle no.3 was chosen as the best nozzle design to use through the remainder of this study. 3.6. Gas Phase Emissions. 3.6.1. Base Point Operation. The conducted tests of flame instabilities and nozzle coking indicate that the bio-oil burner must operate under a narrow range of conditions in order to achieve stable flames and avoid major coke formation on the external surface of the nozzle. Based on many experiments using the new bio-oil 12


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Energy & Fuels

facility, a base operating condition for each fuel batch is determined, under which the exhaust emissions are minimal and the flame is stable. Table 3 summarizes the operating parameters as well as their values at the base point condition for bio-oil batch 2 and batch 3. It should be noted that batch 1 was completely used for initial testing to achieve stable flames of bio-oil.

Table 3. Base point operating condition Parameter

Batch 2

Batch 3

Power input from bio-oil

10 kW

10 kW

Pilot flame power input

0.3 kW

0.3 kW

3.38

3.38

8 SLPM

14.4 SLPM

165°C

165°C

294 SLPM

280 SLPM

0.57

0.58

Swirl number Atomizing air flow rate Primary air preheat temperature Primary air flow rate Equivalence ratio

During each test and after warming up the burner with ethanol combustion, the burner is run on bio-oil at the base point operating condition for 30 minutes until the exhaust flange temperature reaches above 450°C in order to reach a steady state condition. Gaseous emissions are measured to confirm the base point repeatability, which is where both CO and UHC emissions are below their respective detection limits (10 ppm and 3 ppm, respectively) for both bio-oil batches. Average NOx emissions for batch 2 and batch 3 are 106±5 ppm and 124±6 ppm, respectively. Table 4 compares the levels of CO and UHC in the exhaust gases as well as flame stability for both the upgraded and previous burners at their respective base operating conditions. Table 4. Comparison of base point CO and UHC emissions as well as stability between the new burner and the previous one17 Fuel

COa (ppm)

UHCa (ppm)

Stability

New burner

Pure bio-oil

< 10

300

Unstable

New burner

80/20 biooil/EtOH

< 10

0.72) results in flame blowout and hence, increases CO levels due to flame instabilities and poor ignition conditions. 1000

CO (ppm)

800 600 Detached jets Blowout

400 200

No flame

0 0

0.1

0.2

0.3

0.4 0.5 ALR

0.6

0.7

0.8

0.9

Figure 24. CO emissions versus ALR UHC emissions follow a similar trend to CO with respect to atomizing air flow rate since the fuel droplet size plays a key role in the ignition quality and also UHC emissions. As shown in Figure 25, UHC starts to rise at a flow rate below 4.5 SLPM (ΑLR = 0.16) likely due to poor atomization characteristics, which causes the fuel to undergo less thorough burnout. When the air flow rate is reduced to 2.3 SLPM (ΑLR = 0.08), the fuel droplets fails to ignite properly and hence, the flame begins to extinguish. At this flow rate, the fuel droplets dribble out of the nozzle, no longer being atomized and resulting in a sudden rise in unburned hydrocarbon levels in the exhaust gases. If the burner keeps operating at such conditions, combustion would cease with rapidly increasing UHC and CO emissions as the fuel cannot get atomized and ignite. 350 300 250 UHC (ppm)

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200 Detached jets Blowout

150 100 50 No flame 0 0 0.1

0.2

0.3

0.4 0.5 ALR

0.6

0.7

0.8

0.9

Figure 25. UHC emissions versus ALR Contrary to the behavior observed regarding CO emissions at higher atomizing air flow rates, UHC emissions remain below the detection limit even during complete blowout conditions. Under these conditions, the SMD of the fuel spray is at its minimum levels during these tests and the volatile content of the fuel is expected to evaporate rapidly due to the larger surface area of the smaller droplets, inducing greater heat transfer, and oxidizing the fuel into CO rather than UHC emissions. Figure 26 shows that there is an increase in NOx levels when the atomizing air flow rate is increased from 2.3 SLPM (ΑLR = 0.08) up to the base point condition, where the flame is stable and anchored close to the nozzle cap. The reason for the increased NOx emissions is that the fuel droplets undergo more thorough burnout, which along with the improved reactant mixing due to higher air flow rates, favors oxidation of the fuel-bound nitrogen into NOx emissions. Simultaneously however, as the air flow rate is increased, the flame becomes lifted and is anchored further downstream of the nozzle or even blows out as the air flow rate increases above the base point. The higher air flow rate likely slightly deteriorates the ignition quality of the fuel and reduces NOx production, as shown in the plot. As explained previously, NOx emissions with bio-oil are mostly controlled by the conversion of fuel-bound nitrogen and hence, they 19


Energy & Fuels

remain relatively stable throughout most of the atomizing air flow range. The rise in NOx at 21.6 SLPM (ΑLR = 0.77) is probably due to the fluctuations observed in the flame at this condition. 140 120 100 NOx (ppm)

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80 60

Detached jets Blowout

40 No flame

20 0 0

0.1

0.2

0.3

0.4 0.5 ALR

0.6

0.7

0.8

0.9

Figure 26. NOx emissions versus ALR 3.6.6. Pilot Flame Energy. The stoichiometric CH4/O2 flame pilot flame provides a high temperature ignition source for the bio-oil and helps to anchor the flame close to the nozzle. Figure 27 compares the flame at three different pilot energy inputs. After reaching steady combustion at the base point condition in the burner, five other pilot energy conditions were tested to investigate the effects of pilot flame energy on flame stability and pollutant emissions. The flame is found to be stable and anchored to the nozzle between 0.3 kW to 0.15 kW energy input; however, it becomes less stable at an energy input of 0.1 kW (Figure 27.b) and completely unstable when the pilot is turned off (Figure 27.c). Flame (a) in Figure 27 is more desirable for the purposes of this study than flame (b) and flame (c), because it is seated on the nozzle.

(a) 0.3 kW (base)

(b) 0.1 kW

(c) Pilot off

Figure 27. Images of bio-oil flame with varying pilot flame energy inputs For pilot energy inputs below 0.15 kW, the bio-oil flame is unstable close to the nozzle and hence, moved further downstream, burning in the hot regions of the combustor. In terms of CO and UHC emissions, in all cases both pollutants are below the detector threshold, even when the pilot flame is turned off for about 30 minutes. This indicates that bio-oil combustion becomes self-sustaining even without the pilot flame. Although the combustion is self-sustaining at steady state conditions, a pilot energy of 0.3 kW is still necessary for the primary ignition of bio-oil to occur and for the burner to heat up initially. As plotted in Figure 28, there is a positive relationship between NOx emissions and the pilot flame energy. The total energy input of the burner and hence, the burner temperature increases with pilot flame energy, resulting in higher NOx levels in the exhaust of the bio-oil flame. The plot has a small slope because NOx emissions are dominated by the conversion of fuel-bound nitrogen rather than thermal NOx, as previously mentioned.

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Energy & Fuels 140 120

NOx (ppm)

100 80

Minor instabilities

60 40 20 0 0

0.1 0.2 Pilot flame energy (kW)

0.3

Figure 28. NOx emissions versus pilot flame energy 3.6.7. Primary Air and Fuel Preheat Temperature. Through several trials, it was found that having the air heater turned on to maximum (i.e. 1 kW) was necessary for primary ignition to occur. When the flame reached a steady base point condition, the air heater was turned down in order to investigate the effect of primary air temperature on the biooil combustion. CO and UHC emissions remain below their respective detection limits throughout all primary air temperatures. The near complete oxidation of CO and hydrocarbons indicates the efficacy of the refractory linings for the purpose of heat loss reduction. This suggests that there is no need for an alternative heating source. Figure 29 plots the fuel temperature against various primary air preheat temperatures. There is an inverse relationship between fuel temperature and its viscosity; as fuel temperature decreases, the viscosity increases. Moreover, according to equation (1), there is a positive relationship between fuel viscosity and SMD of the spray. As mentioned in section 3.5, spray droplet size has a significant impact on coking tendency of the fuel on the nozzle and therefore, lower preheat temperatures increase the chance of nozzle coking. It is found that the fuel nozzle starts coking at primary air temperatures below 90°C, and the extent of coke formation on the nozzle surface significantly increases at lower preheat temperatures (below 70°C). Although the observed coking does not seem to affect CO and UHC emissions during the course of the experiment, it can potentially deteriorate bio-oil atomization and ignition quality over time during lengthier experiments. Therefore, it is recommended to maintain the air preheat temperature above 90°C in order to avoid coking. 80 Fuel temperature (°C)

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

75 70

Starts coking

65 60 55 50 45 40 25

75 125 Primary air temperature (°C)

175

Figure 29. Fuel preheat temperature versus primary air preheat temperature NOx emissions slightly decrease when the preheat temperature is reduced, as shown in Figure 30. As previously stated, the conversion of fuel-bound nitrogen is the dominant mechanism for producing NOx. However, the slight decrease observed in NOx emissions at lower preheat temperatures is likely the contribution of thermal NOx reduction in the system.

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140 120 100 NOx (ppm)

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

80 Starts coking

60 40 20 0 25

50

75 100 125 150 Primary air temperature (°C)

175

Figure 30. NOx emissions versus primary air preheat temperature 3.6.8. Formaldehyde, Acetaldehyde, and Methane Emissions. Aldehyde emissions are intermediate products formed during the oxidation and combustion of oxygenated biofuels such as bio-oil. Aldehyde and methane emissions have negative impacts on human health and substantially contribute to global warming.43 The FTIR measurements show that formaldehyde (CH2O), acetaldehyde (C2H4O), and methane (CH4) emissions in the exhaust of bio-oil combustion are all below their respective detection limits (10 ppm, 30 ppm, and 10 ppm, respectively) in most cases. This indicates proper oxidation of such intermediate combustion products. However, a detectable formaldehyde emission of around 12 ppm is found when using an atomizing air flow rate of 3 SLPM. Furthermore, methane and formaldehyde emissions of about 30 ppm and 85 ppm, respectively, are found in the lowest atomizing air flow rate case (2.3 SLPM), because of poor fuel atomization and poor ignition quality. 3.7. Particulate Matter Emissions. In addition to the gaseous emissions, the particulate matter emissions in bio-oil exhaust were measured at the base operating point. The amount of carbonaceous residues (CR) and ash in PM emissions were measured via gravimetric analysis of the filters, as explained in section 2.6. Table 7 shows the average values of PM, CR, and ash from the filter deposition measurements. About 90% of the total PM deposited on the filters is constituted of ash, the remainder of which is CR. Table 7. Base point particulate matter emissions

a b

Emissiona

Average valueb (mg/kgfuel)

Standard deviation (mg/kgfuel)

Total PM

504 ±46

14

CR

50

5

Ash

454

12

At an exhaust flow rate of 330 SLPM Four measurements were taken to generate the average values

Table 8 summarizes the CR emissions of pure bio-oil combustion in this study and those of 80/20 bio-oil/EtOH combustion from the previous burner at their respective base operating conditions. It reveals that the new burner provides an adequate temperature to burn out most of the char particles from the fuel. As a result, the final solid phase product exhausted in this study is mostly comprised of incombustible ash from the fuel. Table 8. A comparison of CR emissions between the new burner and the previous one26

Burner setup New burner 26

Previous burner

Fuel

Exhaust temperature (°C)

CR emissions (mg/kgfuel)

As a wt % of total PM

Bio-oil

450-500

50

10

80/20 Bio-

250-300

225

30

22


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oil/EtOH Although PM sampling is performed under isokinetic conditions to obtain representative results, a portion of total PM is inadvertently deposited on the internal surfaces of the burner, which cannot be collected during sampling. Knowing the ash content of fuel, the percentage of PM lost in the combustor is estimated by calculating the amount of ash deposited on the filters as a wt % of fuel mass input. About 50% of the total ash from the fuel is deposited on the internal surfaces of the combustor and exhaust lines and do not reach the filters. This deposition can be explained by the geometry of the combustion chamber, the presence of swirling flow inside the burner, the fuel spray pattern, exhaust pipe cross-sectional reductions prior to the PM sampling line, and the presence of U-joints in the exhaust pipes. 3.7. Heat Exchanger Analysis. The total power input of the burner is provided by three main sources: the fuel input (10 kW), the pilot flame (0.3 kW), and the air heater (1 kW). Using equation (7), the total heat extracted by the exhaust heat exchanger is calculated and the amount of heat loss is then calculated from equation (6). Figure 31 plots the input power of the burner as well as the amount of the heat extracted and lost at different combustion equivalence ratios. Heat is lost from the swirl box, combustion chamber, exhaust pipes, and the heat exchanger unit, with the heat from the exhaust pipes being the most substantial. 12

Power (kW)

10 8 6 4 Extracted heat Total power input Heat loss

2 0 0.45

0.55

0.65 Equivalence ratio

0.75

Figure 31. A thermal energy analysis of the new burner at different equivalence ratios As shown in Figure 31, the amount of heat loss increases when the equivalence ratio is increased. In this case, a lower primary air flow rate results in a higher equivalence ratio and a higher gas temperature. Considering the burner’s thermal inertia, the outer wall temperature can be assumed to remain relatively constant. Therefore, higher gas temperatures create larger thermal gradients, which generate more heat losses to the walls and the room. Figure 32 compares the amount of heat extracted from the new burner and the previous one44 at different equivalence ratios. The power inputs for both the new and previous burners are similar, however, the new burner is found to have a much higher extracted energy as compared to the previous one and so, the new burner has a significantly lower heat loss than the previous uninsulated one as expected. 10 Extracted heat (kW)

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Energy & Fuels

8 6 4 2 0 0.40

New burner (this study) Previous burner 0.50

0.60 0.70 Equivalence ratio

0.80

Figure 32. The amount of heat extracted by the heat exchanger for both the previous and new burners 23


Energy & Fuels

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4. CONCLUSIONS A new refractory lined combustion chamber was designed to replace the previous one in order for the burner to achieve pure bio-oil combustion. It was found that the burner throat had to be modified in order to stabilize the bio-oil flames. Placing a sleeve inside the throat to shrink the area around the nozzle and increase the primary air velocity was the ultimate solution to address initial flame instabilities. Therefore, while it was not possible to stabilize pure bio-oil flames in the previous burner, the new burner can with minimal emissions. When the fuel temperature increases above its minimum boiling point, the bio-oil starts to vaporize within the nozzle cap, resulting in flashing instabilities. It was found that the nozzle needed to be insulated by a ceramic sheet in order to regulate the fuel temperature to prevent these instabilities. Polymerization of bio-oil at high temperatures can result in nozzle coking and clogging. Nozzle no.1 showed a high coking tendency which could not be prevented by adjusting the operating parameters. To solve this problem, three other nozzle designs were tested which decreased the droplet size of the spray. However, there is a compromise between external coking and internal clogging of the nozzle when decreasing its discharge orifice size. Nozzle no.3 was found to have the least coking and clogging tendencies among four nozzles tested in this study. Studies using the previous burner showed relatively high CO and UHC emissions even for flames of 80/20 biooil/EtOH.17,24 However, when using the new burner at the base operating condition, CO and UHC levels are both below their respective detection limits. This demonstrates that the new combustion chamber allows for higher quality combustion by, at least in part, reaching a higher temperature than the previous one, favoring the oxidation of CO and UHCs. The transient operation of the burner showed that when the burner temperature increases, these pollutants tend to decrease until steady combustion is achieved, and CO and UHCs are undetectable in the bio-oil exhaust. Moreover, NOx emissions jump from about 50 ppm to above 110 ppm immediately after switching the fuel from ethanol to biooil, and then remain relatively constant. This indicates that the contribution fuel NOx formation during bio-oil combustion is about 60 ppm. Bio-oil combustion is very sensitive to operating conditions, which in this study have a narrow range due to the limited volatility of the fuel. For instance, an equivalence ratio of around 0.57 is necessary in order to improve mixing between the fuel and air, and to minimize the amount of CO emissions. However, if the primary air flow rate exceeds the base point value, a severe shear is induced in the flame region and the flame temperature decreases, resulting in an increase in CO emissions and a slight decrease in NOx. The equivalence ratio does not affect UHC emissions and they remain below the detection limit in all cases. Moreover, if the swirl number is increased, CO emissions can be minimized at a higher equivalence ratio due to improved fuel/air mixing. It is important to use an optimum atomizing air flow rate at which the flame is anchored to the nozzle and the fuel is properly atomized. Using low atomizing air flow rates results in poor atomization and thus, poor ignition quality of the bio-oil, ultimately increasing UHC and CO emissions. The pilot flame is necessary in order to begin ignition and stabilize the bio-oil flame close to nozzle. In the previous burner, the pilot flame had to be active throughout as turning it off resulted in flame blowout. However, in the new burner, the bio-oil flame becomes self-sustaining once the burner reaches steady state conditions, and even if the pilot flame is turned off at this stage, CO and UHC emissions do not noticeably increase. Using an air heater, the primary air and subsequently the fuel are preheated in order to improve atomization and initiate its ignition. After the bio-oil is ignited and the flame is stabilized, the high gas temperature around the nozzle is enough to maintain the flame, even without using the air heater. However, the fuel temperature would decrease when the primary air temperature is reduced, which likely causes the spray SMD to become larger and coking tendency of the nozzle to increase. With this in mind, a minimum air preheat temperature of 90°C is recommended. Bio-oil has a solid content and tendency to polymerize that results in particulate matter emissions. With the new burner's higher temperature, carbonaceous residues (CR) emissions are reduced. The measurements at the base operating condition show that about 90% of the total PM is composed of ash with the remainder being CR. In the end, the heat exchanger analysis confirms that the new burner has a significantly lower amount of heat loss than the previous one. This observation reinforces the improvement that was seen with regards to flame stability and combustion efficiency of pure bio-oil in the new burner, and the new burner can be easily adapted and scaled for use in industry. Furthermore, since bio-oil burns well in the new burner, there is a reasonable possibility of using this biofuel in small-scale applications, which is an unforeseen, but valuable, experimental outcome. AUTHOR INFORMATION Corresponding Author *Telephone: +1-416-573-1022. Fax: +1-416-978-7753. E-mail: [email protected]. 24


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ACKNOWLEDGMENT The authors would like to thank Professor Hongi Tran and Sue Mao from Pulp and Paper Center at University of Toronto for providing the calorimeter in order to measure HHV of bio-oils. The authors also like to acknowledge BFN and NSERC for being financially supportive of this research. NOMENCLATURE ALR = Atomizing air/liquid fuel mass ratio ASTM = American society of testing and materials CR = Carbonaceous residue(s) CRZ = Central recirculation zone EtOH = Ethanol FID = Flame ionization detector FTIR = Fourier transform infrared spectrometer HHV = Higher heating value HMW = High molecular weight LHV = Lower heating value NOx = Nitrogen oxides (i.e. NO and NO2) OD = Outer diameter PM = Particulate matter ppm = Part per million SLPM = Standard liters per minute SMD = Sauter mean diameter UHC = Unburned hydrocarbon M% = Constant-pressure specific heat capacity d = Diameter 6)'& = Fuel mass flow rate P = Pressure S = Swirl number T = Temperature = Relative air-to-liquid velocity !$'#")*'( = Flow rate of dry gas in PM sampling !,-,#& = Total exhaust flow rate from the burner !"#$%&'( = Actual PM sampling flow rate 012 3 = Molar fraction of water ΔℎJ = Change in enthalpy of water ∆t = Sampling time of the filter ρ = Density σ = Surface tension µ = Dynamic viscosity Φ = Equivalence ratio REFERENCES (1) Steele, P.; Puettmann, M. E.; Penmetsa, V. K.; Cooper, J. E. Forest Products Journal 2012, 62 (4), 326–334. (2) Bridgwater, A. V.; Peacocke, G. V. C. Renewable and sustainable energy reviews 2000, 4 (1), 1–73. (3) Oasmaa, A.; Meier, D. Journal of Analytical and Applied Pyrolysis 2005, 73, 323–334. (4) Gust, S. In Developments in Thermochemical Biomass Conversion; Bridgwater, A. V., Boocock, D. G. B., Eds.; Blackie Academic and Professional: London, U.K., 1997; pp 481−488. (5) Kytö, M.; Martin, P.; Gust, S. In Pyrolysis and Gasification of Biomass and Waste; Bridgwater, A. V., Ed.; CPL Press: Newbury, U.K., 2003; pp 187–190. (6) Wissmiller, D. Pyrolysis oil combustion characteristics and exhaust emissions in a swirl-stabilized combustor. Ph.D. Thesis, Iowa State University, 2010. (7) López Juste, G.; Salvá Monfort, J. J. Biomass Bioenergy 2000, 19, 119–128. (8) Andrews, R.; Patnaik, P. C.; Liu, Q.; Thamburaj R. In Proceedings of Biomass Pyrolysis Oil Properties and Combustion Meeting: Estes Park, CO, September 26–28, 1994; NREL-CP-430-7215, pp 383–391. 25


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