2. ii
Design of a Catalytic Combustor for Pure Methanol and HTPEM
Fuel Cell Anode Waste Gas
Andrew James Bell
Masters of Science in Mechanical Engineering
University of Toronto
2012
Abstract
Transportation sector CO2 emissions contribute to global warming. Methanol generated from clean
energy sources has been proposed as a transportation fuel as an alternative to gasoline or diesel to
reduce emissions. Catalytic methanol-steam reformers can be combined with high temperature
polymer electrolyte membrane (HTPEM) fuel cell systems to create compact electrical power
modules which run on liquid methanol (1). These modules combine the efficiency of a fuel cell system
with the convenience of using a traditional, liquid hydrocarbon fuel.
Catalytic methanol-steam reformers require a heat source as the methanol-steam reforming process
is endothermic. The heat source for this system will initially be from the catalytic combustion of
either pure methanol, during startup, or from HTPEM fuel cell anode waste gas during system
operation. Efficient use of catalyst requires effective premixing of the fuel and air. This study will
investigate parameters affecting premixing and their effect on temperature distributions and
emissions.
3. iii
Acknowledgements
First of all, thanks to Mads Bang and Anders Korsgaard from Serenergy for providing funding and
support for this project while in the midst of building a company from the ground up. Without their
assistance, this project would have never existed. I would also like to thank Leanne Ashworth from
Serenergy for her help with LabVIEW and general enthusiasm about postgraduate work and complex
projects.
From the University of Toronto I would like to thank Professor Murray Thomson for providing
guidance and support when I needed an extra boost (especially during the editing process). He was
there when I needed the help while still giving me the freedom to explore my own ideas. Also from
the University of Toronto, Umer Khan and Coleman Yeung deserve special recognition for their
patience and persistence in assisting with emissions measurements.
From the MIE machine shop, Ryan, Gordon, Fred, Jeff and Terry all provided valuable assistance at
various points in the machining process. Their help was critical in order to meet the various deadlines
for this project. From the administrative staff, Donna Liu and Sheila Baker were extremely helpful. All
of this assistance was greatly appreciated.
A final thank you goes out to the Natural Sciences and Engineering Research Council of Canada
(NSERC) for providing extra funding for the project by means of an Industrial Postgraduate
Scholarship.
4. iv
Table of Contents
Abstract.........................................................................................................................................................ii
Acknowledgements ....................................................................................................................................iii
List of Figures .............................................................................................................................................vii
List of Tables................................................................................................................................................ix
List of Equations...........................................................................................................................................x
Nomenclature.............................................................................................................................................xii
1. Introduction.......................................................................................................................................... 1
1.1. Motivation .................................................................................................................................... 1
1.2. Objectives ..................................................................................................................................... 1
1.3. Overview.......................................................................................................................................2
2. Literature Review................................................................................................................................ 4
2.1. High Temperature Proton Exchange Membrane (HTPEM) Fuel Cells...................................... 4
2.2. Methanol as an Alternative Fuel..................................................................................................7
2.3. Reforming Technologies............................................................................................................. 9
2.3.1. Steam Reforming (SR) ........................................................................................................ 9
2.3.2. Partial Oxidation (POX or CPOX) .......................................................................................10
2.3.3. Autothermal Reforming (ATR) ...........................................................................................11
2.4. Hybrid System Design ................................................................................................................ 12
2.5. Mixing of Non-Reacting Flows...................................................................................................14
2.5.1. Momentum Flux Ratio .......................................................................................................14
2.5.2. Effect of Geometric Parameters on Jet Mixing and Penetration ....................................16
2.5.3. Density Ratio....................................................................................................................... 17
2.5.4. Optimal Mixing Correlations.............................................................................................. 17
2.6. Swirl.............................................................................................................................................19
11. xi
Equation 2-31 - N2O Intermediate Mechanism Step 2 ..............................................................................25
Equation 2-32 - N2O Intermediate Mechanism Step 3..............................................................................25
Equation 2-33 - Unmixedness.....................................................................................................................25
Equation 2-34 – Simplified Local Unmixedness ....................................................................................... 26
Equation 3-1 - Space Velocity......................................................................................................................41
Equation 4-1 - Concentration of Chemical Compound..............................................................................47
Equation 4-2 - Average Cross-Sectional Temperature Difference .......................................................... 50
Equation 5-1 - Final Local Unmixedness Relationship.............................................................................. 54
12. xii
Nomenclature
Symbol Definition
UHC Unburned hydrocarbon
PEM Polymer electrolyte membrane
LTPEM Low-temperature polymer electrolyte membrane
HTPEM High-temperature polymer electrolyte membrane
MEA Membrane electrode assembly
GDL Gas diffusion layer
PBI Polybenzimidazole
ZEV Zero emission vehicle
NFPA National Fire Protection Association (USA)
SR Steam reforming
POX Partial oxidation
CPOX Catalytic partial oxidation
ATR Autothermal reforming
J Momentum flux ratio
Jopt Optimal momentum flux ratio
ρm Cross-flow density
ρj Jet density
Vj Jet velocity
Um Cross-flow velocity
13. xiii
Am Cross flow area
Aj Jet area
mj Total mass flow rate of jets
mm Total mass flow rate of cross-flow
ppm Parts per million
H Duct height
s Jet spacing
d Jet diameter
D Diameter of cylindrical cross-flow
C Dimensionless parameter denoting unmixedness in Holdeman correlation
R1/2 Mid-area radius
n Number of jets (Section 2.5) or number of swirl blocks (Section 2.6)
R Swirl generator exit radius
Rh Swirl generator inner radius
B Depth of swirl blocks
α Fixed swirl block angle
ξ Adjustable swirl block angle
ξm Maximum opening angle
S Dimensionless parameter representing swirl in the cross flow
UM Unmixedness
σ Standard deviation of total carbon molar fraction
14. xiv
μ Mean value of total carbon molar fraction
AWG Anode waste gas
DAQ Data acquisition
FTIR Fourier transform infrared spectrometer
FID Flame ionization detector
RSME Root mean squared error
SLPM Standard litres per minute
15. i
1. Introduction
1.1. Motivation
Transportation sector CO2 emissions contribute significantly to global warming. Methanol generated
from clean energy sources has been proposed as a transportation fuel as an alternative to gasoline
or diesel. Liquid hydrocarbon fuels are easy to store and transport however internal combustion
engines are not typically as efficient as hydrogen fuel cells. Catalytic methanol-steam reformers can
be combined with high temperature polymer electrolyte membrane (HTPEM) fuel cell systems to
create compact electrical power modules which run on liquid methanol (1). These modules can be
used for a wide variety of applications from range extenders in electric cars to backup power for
telecommunications.
HTPEM fuel cells with integrated methanol-steam reformers combine the efficiency of a fuel cell
system with the convenience of being able to use a traditional, liquid hydrocarbon fuel. Producing
hydrogen on demand from these systems has been suggested as a cost feasible method for reducing
transport sector CO2 emissions.
Catalytic methanol-steam reformers require a heat source to operate as the methanol-steam
reforming process is endothermic. The heat source for the Serenergy system will initially be from the
catalytic combustion of either pure methanol, during startup, or from HTPEM fuel cell anode waste
gas during system operation. Utilizing the catalytic combustion of pure methanol on startup will
eliminate the need to rely on electrical power to heat both the reformer and fuel cells up to the
desired operating temperature.
1.2. Objectives
Efficient use of catalyst requires effective premixing of the fuel and air. The purpose of this study is
to analyze specific parameters affecting premixing, and their effect on temperature distributions and
emissions in the catalytic combustion process. This will help to determine which design parameters
are critical for minimizing the production of carbon monoxide and formaldehyde while maintaining
minimal hydrocarbon slip. Furthermore, this study provides empirical evidence to aid in the
certification process of future methanol-steam reformer/HTPEM fuel cell hybrid systems from
Serenergy.
1
16. 2
This work will build on previous two previously published studies. The first was completed by
Ditaranto et al. in 2007 and was called Experiments in a Catalytic Reactor Burning the Anode Off-Gas of
a Methanol Fuel Cell (2). The second was completed by Wiese et al. in 1999 and called Emission
Behaviour of a Catalytic Burner Fuelled with Mixtures of Hydrogen and Methanol (3). In the previous
investigations, a variety of operating conditions and fuel compositions were analyzed and emissions
data was collected. Specifically, temperature and emissions profiles were created in order to
investigate the effect of altering parameters such as the composition of the anode waste gas and
the amount of excess air among others. Some unique features of this investigation include:
1. A new, proprietary catalyst composition
2. Non-preheated air
3. A unique anode waste gas composition
4. A larger scale of burner
5. A unique burner design
Specifically, the emissions from two different fuels will be analyzed.
1. Pure methanol (evaporated prior to combustion)
2. A simulated anode waste gas mixture consisting of H2, CO and CO2
The goal of this study is to design and build a catalytic burner with flexible operating parameters,
such that parameters can be altered and analyzed in order to determine those which are critical for
minimizing CO, UHC and formaldehyde emissions.
1.3. Overview
This thesis is divided into the following sections.
1. Introduction
2. Literature Review
3. Design
4. Experimental Methodology
5. Results and Discussion
6. Conclusions
17. 3
The Introduction and Literature Review sections provide an overview of research activity in the areas
of catalytic combustion, mixing of gases and previous experiments with similar objectives. The
Design section outlines decisions that were made during the development of the burner. The
Experimental Methodology section provides a detailed overview of the experiments completed. It
also provides a summary of the calibration techniques for the various emissions testing devices and
how they were subsequently utilized during this study. The Results and Discussion section includes
relevant sensitivity studies and a presentation and analysis of the emissions results. The conclusions
are presented in the final section along with suggestions for future research work in this area.
18. 4
2. Literature Review
Catalytic methanol-steam reformers require a heat source to operate as the methanol-steam
reforming process is endothermic. The heat source for the Serenergy system will initially be from the
catalytic combustion of either pure methanol, during startup, or from HTPEM fuel cell anode waste
gas during system operation. Utilizing the catalytic combustion of pure methanol on startup will
eliminate the need to rely on electrical power to heat both the reformer and fuel cells up to the
desired operating temperature.
In order to provide context for the design requirements of the catalytic burner, a better
understanding of the system components is required. This review covers the basic theories and
operational guidelines for PEM fuel cells and reformers before reviewing papers on mixing of non-
reacting cross flows and finally, previous similar experiments on catalytic combustion.
2.1. High Temperature Proton Exchange Membrane (HTPEM) Fuel Cells
In a typical proton exchange membrane (PEM) fuel cell, a proton conducting polymer membrane
separates the anode and cathode. See Figure 1 below for the layout of a typical PEM fuel cell.
Figure 1 - Typical PEM Fuel Cell (4)
In this type of fuel cell, hydrogen flows into the anode and is split into hydrogen ions (protons) and
electrons. Since the membrane only conducts protons, the loose electrons can be collected and used
to power a device. Once the hydrogen ions permeate through the membrane, they combine with the
oxygen in the cathode air stream as well as the loose electrons to form water. Heat is created in this
process as well and either exits through the cathode stream or diffuses through the fuel cell stack.
The reactions for this process are as follows.
19. 5
2 · 4 ·
4 ·
Equation 2-1 – Anode Reaction
4 ·
4 · 2 ·
Equation 2-2 - Cathode Reaction
2 · 2 ·
Equation 2-3 - Overall Reaction
The theoretical maximum open circuit voltage which can be achieved from a single cell fuel cell using
these reactions is 1.16V at 80˚C and 1atm (5). This cell voltage is not achievable under load because as
current is drawn, the voltage drops. In order to achieve the higher voltages necessary in many
applications, several cells are commonly joined together in series to form an assembly called a stack.
While Figure 1 shows a simplified fuel cell assembly, in reality the membrane layer depicted in the
center is not a single layered component. The term ‘Membrane’ in the diagram actually represents a
membrane electrode assembly (MEA) consisting of several layers. See Figure 2 below for an
exploded view of the MEA.
Figure 2 - Membrane Electrode Assembly (6)
Each of these layers performs a very specific task. The gaskets prevent gas from bypassing the
membrane. The gas diffusion layer (GDL) is typically joined with the electrode layer to form a gas
diffusion electrode, which is used to split the hydrogen atom (anode side) and the oxygen atom
(cathode side). Splitting the hydrogen atom is relatively simple process, completed with a platinum
catalyst which is embedded in the GDL. A catalyst is required to accelerate the reactions, which are
slow due to the low temperatures PEM fuel cells operate at. Splitting the oxygen atom is more
energy intensive and accounts for much of the electrical losses in a PEM fuel cell. The cathode side
GDL uses a platinum catalyst as well, mainly due to lack of superior commercially available options.
20. 6
PEM fuel cells have other unique properties which make them ideal for use in automotive
applications. They can be started rapidly due to their low operating temperature and can adapt very
quickly to fluctuations in load. Additionally, they also can be packaged to provide a high power
density and are not sensitive to orientation offering excellent design flexibility. Furthermore they are
efficient, operating at 40-60% of maximum theoretical voltage in most applications.
Low temperature PEM (LTPEM) fuel cells typically operate at a temperature of 80˚C, which is limited
by the properties of the available membrane materials. Most LTPEM fuel cells use a DuPont material
called Nafion for the membrane layer. This material relies on liquid water humidification to transport
the protons. As the fuel cell approaches 100˚C, the membrane dries out and the fuel cell stops
generating current thus creating an operating temperature limit.
Recently, a new type of membrane has been created which uses either phosphoric acid or
polybenzimidazole (PBI) in place of water, effectively eliminating all water management issues. Fuel
cells using these new membranes are referred to as high temperature PEM (HTPEM) fuel cells. These
membranes allow operating temperatures of up to 220˚C. Some added benefits of using high
temperature membranes include
- Increased resistance to carbon monoxide poisoning (up to 3% CO in anode gas supply versus
20ppm maximum in LTPEM FC’s)
- Potentially higher efficiency
- Higher power density
- Easier cooling (greater allowable temperature difference versus ambient air)
- Easier controllability (no water management issues)
A further benefit of HTPEM fuel cells is the ability to use untreated reformate gas directly in the
anode. This is a major advantage from a commercial application perspective. The major disadvantage
of using LTPEM fuel cells in automotive applications is that an on-board hydrogen supply is required.
Due the purity requirements of LTPEM fuel cells (20ppm CO for example), it is not practical to
produce hydrogen on-demand for these systems. HTPEM fuel cells have much more lenient purity
requirements. As a result, a low-cost methanol steam reformer can be used to produce hydrogen on
demand, with the entire system able to be easily packaged and mounted in a vehicle. The end result
allows for the driver to refuel the vehicle with a liquid instead of a gas. This development increases
the safety of the system and eliminates a critical impediment to mass adoption.
21. 7
One of the last remaining barriers to the mass adoption of both HTPEM and LTPEM fuel cells is the
relatively high cost of membrane materials and catalysts. The concentration of precious metals in the
latest generation of PEM fuels cell MEA’s has been greatly reduced however and economies of scale
should make mass production possible when the market demands it. Companies such as General
Motors, Honda and Toyota have all contributed greatly to this aspect of MEA development.
2.2. Methanol as an Alternative Fuel
One of the major barriers to the mass adoption of fuel cell systems in transport applications has been
the lack of an appropriate energy storage medium. Gaseous hydrogen is inconvenient to store and
must be highly compressed to create the energy densities required for long range travel, creating an
unnecessary danger for motor vehicle operators. In liquid form, the hydrogen must be kept
extremely cold to avoid evaporation – an inconvenience especially in equatorial regions.
Compressing or liquefying hydrogen also wastes a significant amount of energy in itself, further
decreasing the overall efficiency of these systems. Other technologies exist for hydrogen storage
however none have proven cost effective or user friendly enough for public use.
An alternative approach would be to use the hydrogen contained in liquid hydrocarbons as a fuel.
Methanol is an ideal choice for this application due to its high hydrogen to carbon ratio and low
reforming temperature using modern Cu/ZnO/Al2O3 catalysts. Also, its miscibility in water is a huge
advantage in the steam reforming process. While the arguments for and against using methanol as
fuel are numerous, this paper will simply highlight some of positives and negatives of the fuel.
One of the most important concerns when selecting a fuel is the potential environmental impact of
an uncontrolled release, should such an accident occur. While it is true that a large release of
methanol into the ground water, surface water or soil does have potential to adversely affect the
environment, methanol has a significantly shorter half-life than gasoline once spilled, making it a
safer selection yet still one with potential risk. See Table 1 below for a comparison of half-lives with
gasoline components.
22. 8
Environmental Medium
Methanol Half Life
(days)
Benzene Half Life
(days)
Iso-Octane Half
Life (days)
Soil (based on un-acclimated grab
sample of aerobic/water suspension
from groundwater aquifers)
1-7 5-16 3-15
Air (based on photo-oxidation half
life)
3-30 2-20 4-10
Surface Water (Based on un-
acclimated aqueous aerobic
biodegradation)
1-7 5-16 5-14
Groundwater (based on un-
acclimated grab sample of
aerobic/water suspension from
groundwater aquifers)
1-7 10-730 NA
Table 1 - Half Lives of Methanol, Benzene and Iso-Octane (Gasoline) (7)
In the event of a catastrophic methanol spill, the methanol will rapidly dilute to low concentrations
followed by a rapid subsequent biodegradation. If clean-up measures are required, they must be
implemented much more quickly than in a petroleum spill in order to capture the fuel before
significant dilution. The natural cleanup times for methanol in general are faster than active cleanup
times for petroleum spills making it a much safer choice overall.
In terms of fire safety, methanol is a Class IB fuel. This is the same flame class as gasoline according
to the United States National Fire Protection Association (NFPA) (7). While methanol is similarly
volatile in comparison with gasoline, it has a lower vapour density meaning that it tends to disperse
more quickly to non-combustible concentrations. Furthermore, methanol vapour must be four times
more concentrated than gasoline vapour to form a combustible mixture in air making it significantly
less dangerous. The main disadvantage of a methanol flame is that it is invisible to the naked eye.
This makes rescue in the case of an accident more dangerous (7). Another important disadvantage is
that methanol can potentially form combustible mixtures in the headspace of a fuel tank.
Currently, the majority of methanol is produced from synthesis gas, which is obtained through the
catalytic reforming of fossil fuels. Traditionally methanol was prepared through pyrolysis of wood. As
long as natural gas remains abundant, direct oxidative conversion of methane into methanol could
be the best short-term source as noted by Nobel Laureate George Olah in Beyond Oil and Gas: The
Methanol Economy (8). Essentially any material that can be gasified can be catalytically transformed
into methanol. Enerkem, a Canadian company based out of Quebec, uses municipal solid waste as a
feed stock for example (9).
23. 9
2.3. Reforming Technologies
There are many different methods of obtaining hydrogen from hydrocarbons used in industry. While
steam reforming is the most common, there are other options. The top 3 most popular methods are
1. Steam reforming (SR)
2. Partial oxidation or catalytic partial oxidation (POX or CPOX)
3. Autothermal reforming (ATR)
All three of these methods can be packaged and used for providing an on-board hydrogen source
from a liquid fuel. A carbon monoxide (CO) removal stage between the reformer and the fuel cell is
often required however as PEM fuel cell membranes are sensitive to CO poisoning. HTPEM fuel cells
raise the acceptable CO limit from several ppm to approximately 3% (5) which is advantageous.
There are several major differences between industrial reforming and on-board reforming when
considering the operational requirements. On-board reformers are orders of magnitude smaller in
scale and have unique packaging constraints. For use on a passenger vehicle, the reforming system
must be very compact to prevent it from impeding on passenger or storage space. Furthermore,
industrial reformers operate on a vastly different duty cycle than that required for a passenger
vehicle, as start-up and shut-down procedures for industrial reformers do not typically consume a
large percentage of total operating time. As well, the lengths of these procedures are not critical for
performance. Finally, once started, industrial reformers tend to operate very near to steady state
conditions. In comparison, an on-board passenger vehicle fuel reforming system must start up very
quickly, have the ability to operate at a variety of potentially varying load points depending on the
operating strategy and, most importantly, be very compact.
2.3.1.Steam Reforming (SR)
In a steam reformer, steam containing a hydrocarbon fuel reacts in the presence of a catalyst to
produce a synthesis gas containing hydrogen. In methanol-steam reformers, which generally contain
the most commonly available Cu/ZnO/Al2O3 catalysts, three main reactions occur (10).
3 · 49.4/
Equation 2-4 – Methanol-Steam Reforming Reaction #1
2 · 90.4/
Equation 2-5 – Methanol-Steam Reforming Reaction #2
24. 10
41.4/
Equation 2-6 – Methanol-Steam Reforming Reaction #3
Equation 2-4 is the primary methanol-steam reforming reaction. It is the dominant producer of
hydrogen as well as the rate limiting step in the process. Equation 2-5 is the methanol decomposition
reaction. It is less favourable under typical operating conditions (200-300C, 1-10bar) although it is the
dominant producer of CO. Equation 2-6 is the water gas shift reaction. This reaction has the potential
to consume the CO produced in Equation 2-5 however the reaction rate is lower than Equation 2-4
under typical operating conditions, thus residence time must be longer to reduce CO emissions.
Temperature profiles across the reactor bed as well as the steam-to-carbon ratio of the fuel
(adjusted by varying the amount of water added to the methanol) can both be manipulated to
influence the kinetics of the water gas shift reaction (Equation 2-6) in order to minimize the CO
produced (11).
It is important to note that the percentage of water content is a crucial aspect of the methanol-
steam reforming reaction. Water affects selectivity towards hydrogen if it is below the stoichiometric
ratio. Experimentally, it has been determined that steam to methanol ratio of 1.5-2.0 is ideal (12).
Furthermore, while coke (pure solid carbon) is very unlikely to form, the prospect is increased when
water is present below the stoichiometric ratio through the following two reactions:
Equation 2-7 - Coking Reaction #1
Equation 2-8 - Coking Reaction #2
The methanol-steam reforming process is endothermic as shown in reaction equations listed above.
This greatly increases the system safety as the risk of runaway reactions is minimal. Most compact
fuel cell systems with integrated reformers use steam reformers.
2.3.2. Partial Oxidation (POX or CPOX)
Partial oxidation (POX) reforming, in general terms, is the incomplete combustion of a fuel. Fuel
reacts with a below stoichiometric amount of oxygen (usually in the form of air) to form a synthesis
gas consisting of CO and hydrogen. An important design consideration when using a POX reactor is
the highly exothermic reaction taking place, and thus a cooling system is mandatory. With careful
design, this heat source can be used elsewhere in the system. A catalytic partial oxidation reactor
25. 11
(CPOX) uses a catalyst to lower the temperature. Due to the high temperatures involved and the
presence of nitrogen in air, it is possible to form NOx and ammonia compounds – a significant
disadvantage of this method. This may require treatment as both of these products could potentially
harm the fuel cell MEA’s.
In POX and CPOX reactors, it is possible to vary the temperature, pressure and oxygen to carbon
ratio (O/C). The effect of altering these parameters is briefly discussed below (5).
- Higher temperatures increase the ratio of hydrogen and CO in the exhaust while burning off
carbon particulate matter.
- Higher pressures increase the concentration of ammonia and solid carbon particles in the
reformate stream and decreases overall efficiency.
- Increasing the O/C ratio displaces the formation of solid carbon particles to lower
temperatures but directly decreases the hydrogen concentration in the reformate stream.
CO concentration is also reduced.
This type of reactor is relatively easy to build but requires careful fuel control to prevent runaway
reactions, very good insulation to protect other system components and an intermediate reformate
cleanup stage. Depending on the design of the system, the reformate stream may also have to be
cooled prior to entering the fuel cell in order to prevent the MEA’s from melting.
2.3.3. Autothermal Reforming (ATR)
Autothermal reforming (ATR) can be defined as a combination of steam reforming with a partial
oxidation reactor (5). ATR occurs in the presence of a catalyst that controls the reaction pathways
which then determines the relative proportion of SR and POX reactions. The SR reaction absorbs
some of the heat generated by the POX reaction, which then limits the overall reactor temperature.
A CPOX reaction is required however to limit the temperature to a range conducive to steam
reforming. A mildly exothermic process is the end result.
ATR fuel processors are a reasonable compromise between POX and SR systems. They start up and
respond more quickly than a SR system while operating at a lower temperature than a POX reactor.
The overall efficiency and hydrogen concentration in the reformate stream are both high, although
these values are strongly affected by system design. CO concentration in the reformate stream tends
26. 12
to be higher than in SR systems but lower than POX systems. Some general observations about the
operation of an ATR system are presented below (5).
- Increasing the pressure decreases hydrogen concentration in the reformate stream
- Increasing the temperature increases both the concentration of hydrogen and CO in the
reformate stream.
- Increasing the S/C ratio beyond the stoichiometric condition lowers the concentration of CO
in the reformate
- Increasing the O/C ratio lowers the hydrogen concentration and raises the temperature
While higher efficiencies are possible with ATR systems, they tend to be more difficult to control and
more complex to build. It is the system designer’s responsibility to decide whether the extra
complexities are warranted.
2.4.Hybrid System Design
Catalytic methanol-steam reformers can be combined with high temperature polymer electrolyte
membrane (HTPEM) fuel cell systems to create compact electrical power modules (1). A typical
block diagram layout for these modules is shown below in Figure 3. This figure shows both startup
conditions and steady state operation modes.
Figure 3 - HTPEM/SR Block Diagram
27. 13
These modules can be used in any application which requires electricity - anything from range
extenders in electric cars to residential back-up power. HTPEM modules with integrated methanol
steam reformers are most efficiently utilized in applications which do not require sudden load
changes. As a result, they are often coupled with battery packs to handle peak loads. A DC/DC
converter is often mounted between the fuel cell and the battery pack to further minimize the
fluctuations in voltage that the fuel cell experiences. A typical system layout for a vehicle is shown
below in Figure 4.
Figure 4 – HTPEM/SR System Layout
By using a battery pack in combination with the fuel cell module, many obstacles are overcome.
Quick startup becomes less critical because the charging and driving functions of the vehicle are
separated. The fuel cell can be easily controlled as the voltage is not tied to pack voltage but instead
to the DC/DC converter. Since the fuel cell can operate at roughly steady state, the methanol-steam
reformer is able to operate more efficiently. The HTPEM/SR system output is chosen such that the
average power requirements of the vehicle can be met. This value is surprising low under most
driving conditions (usually 20kW based on an internal Serenergy study). In designing the system
output this way, the vehicle range is only limited by fuel capacity, as in a standard gasoline or diesel
vehicle. A further advantage to using an HTPEM/SR system in a vehicle is that any excess heat from
the system can be recycled for controlling the cabin temperature, increasing efficiency.
28. 14
Emissions advantages are significant. No NOx is present in the exhaust and no particulate matter is
formed. Carbon dioxide output is minimized due to the increased efficiency. These are all
improvements over a gasoline or diesel powered internal combustion engine.
2.5. Mixing of Non-Reacting Flows
Many studies have been conducted on the mixing of non-reacting flows as there are numerous
industrial applications which involve this process. The interaction between jets and a non-reacting
cross flow involves relatively complex fluid dynamics relationships which are discussed below.
2.5.1.Momentum Flux Ratio
The momentum flux ratio J was determined to be an important flow parameter affecting mixing
according to a study by Holdeman in 1997 (13). J is defined below in Equation 2-9 as the momentum
of the jets to the momentum of the cross-flow air stream.
! · #
$/ !% · %
$
Equation 2-9 - Momentum Flux Ratio
In Equation 2-9, ρj and Vj refer to the density and velocity of the jet while ρm and Vm refer to the
density and velocity of the main flow respectively. Geometric parameters such as jet diameter and jet
spacing were investigated as well as the effects of cross-flow turbulence level, momentum flux ratio,
Reynolds number, jet to cross flow density and jet to cross flow velocity ratio on mixing efficiency.
Geometric parameters and density ratios will be discussed in following sections.
An important summary article by Holdeman was referenced in this section (14). Holdeman reports on
the major findings in the field of mixing of jets in a subsonic confined non-reacting cross-flow
through analysis of many experimental and computational findings, particularly from NASA
supported investigations. The studies were primarily simulations of the flow in the dilution zone of
the combustion chamber in a gas turbine engine. The objectives were to identify dilution zone
configurations that provide a desired mixing pattern within a specified combustor length. The
optimal pattern is one that gives the lowest amount of unmixedness (therefore the best mixing)
over a minimum downstream distance. The studies covered the following ranges:
- Density ratio ρj/ ρm : 0.5-2.2
- Downstream distance (measured in multiples of duct height) x/H : 0-2
29. 15
- Momentum flux ratio J : 5-105
- Area ratio Aj/Am : 0.025-0.1
- Mass flow rate ratio (total jet mass flow rate over total cross-flow mass flow rate) mj/mm :
0.075-0.36
The common variables analyzed across the studies were the momentum flux ratio, the jet to cross-
flow density ratio, jet spacing s and jet diameter d. As expected, the mixing is generally found to
improve with increasing downstream distance. The most crucial factor affecting mixing was
determined to be the momentum flux ratio J. There is an optimum J value, JOpt, which produced the
most efficient mixing for each configuration analyzed. Any deviation from JOpt yielded a much longer
downstream distance to achieve similar mixing performance.
Since jet penetration largely depends on the J value for any given jet configuration, three situations
can arise.
- Under penetration
- Optimum penetration
- Over penetration
Under penetration occurs when J is too low. The jet fluid is trapped along the walls of the cross-flow
pipe. When J equals JOpt, the centerline of the jet reaches the radial center point of the cross-flow
pipe. If J is creased past JOpt, the jets impinge on the center of the duct. If J is much higher than JOpt,
this impingement can lead to back flow. A visual representation of these scenarios is shown below in
Figure 5.
30. 16
Figure 5 - Jet Penetration Depending on Value of J (15)
Of further note, it was determined that mixing under the experimental conditions listed earlier is not
significantly affected by the cross-flow Reynolds number or the turbulence intensity. Rather, the
amount of swirl in the cross flow was determined to have a major effect (16). Cross-flow swirl
improves mixing and, due to the lateral momentum transfer which causes the jets to shift away from
normal to the pipe surface, reduces jet penetration. When swirl is present, JOpt is higher than when
swirl is not present.
2.5.2. Effect of Geometric Parameters on Jet Mixing and Penetration
When the number of jets is constant and the momentum flux ratio J is maintained by adjusting the
mass flow ratio, the diameter of the jets does not affect mixing or jet penetration (13). Reducing the
number of jets does create a measurable increase in jet penetration according to the same study.
According to a previous study by the same author, increasing the jet spacing around the
circumference s, at constant J and duct height H increases penetration at the cost of lateral
uniformity (14). For a constant area ratio Aj/Am (ratio of jet area to cross flow area), jets with higher
jet individual area tend to over-penetrate while jets with a lower individual area tend to under-
penetrate. Optimum mixing occurs when the jet stream penetrates 50-65% of the way to the duct
centerline at a distance of one duct radius downstream of the entry point (17). Two other general
observations of note are that as the number of jets increases, so does the optimum J value JOpt and
larger jet spacing requires a greater downstream distance to achieve optimal mixing (18).
31. 17
When considering a setup with inline jets, it was observed that at x/H = 0.25, the most efficient
mixing occurs when the jets penetrate to the center of the mixer height without impinging on one
another (16). If J is above or below the optimal value, over or under penetration occurs respectively
which results in poor mixing behaviour. In the same study, jet diameter was discovered to have a
non-negligible effect on mixing which contradicts the results from the study by Holdeman in 1993.
All references do agree however on the effects of penetration and jet spacing on the quality of
mixing. Jet diameter can have an effect on mixing depending on the range of momentum flux ratios.
A later section on optimum mixing correlations examines this effect.
2.5.3. Density Ratio
Assuming a constant momentum flux ratio, the effect of the density ratio appears to be negligible.
Density ratio has a second order influence on profile shape and jet penetration according to
Holdeman. The density ratio is incorporated into the momentum flux ratio anyways, thus accounting
for any variation in this ratio (14).
2.5.4. Optimal Mixing Correlations
There are numerous correlations presented in literature based on experimental and computational
models however Holdeman’s correlations are the most frequently reported. According to Holdeman,
an inversely proportional relationship between the momentum flux ratio J and the relative spacing
s/H show similar temperature distributions over a range of J. When the spacing is increased, causing
the momentum flux ratio to decrease, the circumferential unmixedness increases. Jet penetration
and center-plane profiles are comparable when the following relationship is satisfied (14).
'/$ · (
Equation 2-10 - Holdeman Correlation
In the above correlation, C is a constant. There are three flow regimes depending on the range of
values for C. The following regimes are based on experimental and empirical data for single sided
injection in a rectangular cross-flow.
- C 1.5 : Jets under-penetrate and jet fluid remains trapped along duct wall
- C = 2.5 : Mixing is obtained over the minimum downstream distance
- C 5.0 : Jets over-penetrate and jet fluid is concentrated in the center of the duct
32. 18
With respect to jet spacing, the optimum momentum flux ratio JOpt is determined using Equation 2-11.
This equation is derived by solving Equation 2-10 for J and substituting a value of 2.5 for C.
)*+ 6.25 · /'$
Equation 2-11 - Holdeman JOpt for Rectangular Duct with One-Sided Injection
In a subsequent paper, Holdeman stated that for a cylindrical duct, the radius R corresponds to the
duct height H assuming one-sided injection in a rectangular duct (13). For a cylindrical duct, the
optimum jet spacing was specified at the radius. This divides the cylindrical duct into equal areas. The
relationship between the spacing of the jet centerlines to the number of holes around the
circumference of the duct is given below in Equation 2-12 (13).
'
2 · . · /0/
1
Equation 2-12 - Holdeman Jet Spacing in a Cylindrical Duct
Where R1/2 is defined as:
/0/
√2
Equation 2-13 - Definition of R1/2
Substituting Equation 2-13 into Equation 2-12, and substituting the resulting s/H into Equation 2-10,
results in an approximation for the number of round jets in a circular duct.
1
. · (2 ·
Equation 2-14 - Holdeman Jet Quantity Approximation
Solving Equation 2-14 for J and substituting in a value of 2.5 for C results in an equation for JOpt in
terms of the number of jets n.
)*+ 0.32 · 1
Equation 2-15 - Holdeman JOpt Approximation for Cylindrical Duct
A study by Hatch et al. also determined that J is the most important parameter affecting the flow
field (19). In this study, the authors used an apparatus containing a cylindrical duct with 8, equally
spaced, circumferential jets. The flow fields for momentum flux ratio values of 26.7, 55.4 and 84.2
33. 19
were examined. Equation 2-15 predicts a JOpt value of 20.48. The flow field when the momentum flux
ratio was 26.7 yielded the most efficient mixing as J was closest to JOpt.
There are numerous other empirical relationships for different jet geometries (slots, square holes,
etc.) however round holes are the most flexible from a prototype manufacturing perspective.
Equation 2-15 and Equation 2-11 will be used as design guidelines for the experimental apparatus in
this investigation.
2.6.Swirl
In the research papers discussing mixing of gases in non-reacting cross-flows, there were numerous
references to the concept of swirl. In particular, it was noted that increasing the non-dimensional
swirl number S in cylindrical ducts increases the optimal value of the momentum flux ratio JOpt (13). A
swirling flow is one that contains an angular momentum component in addition to an axial
momentum component. If the ratio of angular to axial momentum increases past a certain threshold,
the large radial pressure gradients induced by the swirl can cause the formation of a low pressure
zone in the central region of the flow. The decay of these radial pressure gradients downstream can
cause a strong axial pressure gradient opposite the bulk flow. Fluid can ultimately be forced back
towards the upstream low pressure region due to this reverse axial pressure gradient. This
interaction can create a central recirculation zone which can be advantageous in combustion
applications (20).
If the swirl is strong enough, the central recirculation zone that forms draws hot exhaust products
towards the base of the flame (20). This helps with ignition and flame stability and can lead to a
reduction in UHC, CO and particulate matter emissions. A significant amount of turbulence is also
generated in swirling flows which promotes mixing between the fuel and oxidizer (20). This can
extend the lean blow-out limit and shorten the length of the flame. The use of swirling flows is an
essential design feature in most industrial burner systems.
Both the axial flux of axial momentum Gx and the axial flux of angular momentum Gф are conserved
in free swirling jets (21) (22).
34 5 !2.676 5 82.676 91' 1
:
;
:
;
Equation 2-16 - Axial Flux of Axial Momentum
34. 20
3 5 =6$!2.676 91' 1
:
;
Equation 2-17 - Axial Flux of Angular Momentum
In the above equations, U, W and p represent the axial and tangential components of velocity and
the static pressure, respectively, along any cross-sectional plane in the flow field. Using this
information, swirling flows can be characterized by the non-dimensional swirl number S.
3?
34/
Equation 2-18 - Non-Dimensional Swirl Number
S is a non-dimensional quantity which represents the ratio between angular and axial momentum
fluxes. It provides a method for comparing the swirl intensities of various flows. In Equation 2-18, R
represents the radius of the plane by which the swirling flow exits. This is more clearly illustrated by
Figure 7 in the following section.
2.6.1.Moveable Block Swirl Generator
The final design iteration, which will be discussed in a later section, is shown below in Figure 6. It is
very loosely based on a bio-oil swirl burner design by Tzanetakis et al (22).
35. 21
Figure 6 - Moveable Block Swirl Generator
Room temperature combustion air enters the air box at which point it is forced through either a
radial or a tangential channel in the swirl generator. This design has the ability to impart a varying
degree of angular momentum to the flow by varying the opening size of the radial and tangential air
passages.
The ability to continuously vary the amount of swirl present during catalytic burner operation is an
important design feature of this catalytic burner. Adjustment is accomplished by varying the angle
between the radial and tangential inlets (ξ). When the tangential inlets are closed, all of the air
passes through the radial inlets and into the burner with zero angular motion. This represents the
zero swirl case (S=0). When the radial passages are closed, the incoming air is imparted with the
maximum swirl possible given the geometry of the assembly. It should be noted that the pressure
drop across this type of swirl generator is also lower than for other types of swirl generators and
turbulence generators in general (22). Figure 7 below shows a conceptual top view diagram of a
moveable block swirl generator.
36. 22
Figure 7 - Geometry of a Moveable Block Type Swirl Generator (22)
Descriptions of the various design parameters have been summarized in Table 2 below.
Parameter Description
n Number of swirl blocks
R Swirl generator exit radius
Rh Swirl generator inner radius
B Depth of swirl blocks
α Fixed swirl block angle
ξ Adjustable swirl block angle
ξm Maximum opening angle
Table 2- Parameter Descriptions for Moveable Block Swirl Generator
The swirl number at the exit plane of the swirl blocks (position noted on Figure 6) can be estimated
using Equation 2-19, a theoretical expression developed for moveable block geometries (23) (24).
3?
34/
~ A
/
2B
C1 D
/E
/
F
G
Equation 2-19 - Swirl Number
The dimensionless parameter A in the above equation is represented by Equation 2-20.
37. 23
A
2.
1H%
sin L
cos L O1 tan L tan
H
2
R D
H
H%
F
S1 O1 cos L T1 tan L tan
H
2
UR D
H
H%
FV
Equation 2-20 - Non-Dimensional Parameter 'A' for Swirl Number
It is important to note that the expression for estimating S presented by Leuckel (23) is not the same
as what is reported by Beér and Chigier (21). Of further note, Fudihara et al. (24) states that the
expression used by Beér and Chigier is missing a squared term in the denominator. According to an
analysis by Tzanetakis et al. (22), similar swirl numbers are predicted across all of the equations with
the largest discrepancy being approximately 17% at maximum swirl. The corrected Beér and Chigier
expression shown in Equation 2-19 was used to estimate the swirl for this investigation (22). Given
the potential sources of error associated with the theoretical calculation of S, any of the equations
would be valid although only at the plane of interest.
2.7. Formation of Oxides of Nitrogen
When combusting fuels that contain no nitrogen, there are three possible chemical reaction
mechanisms that can form nitric oxide (NO) from air.
- Thermal or Zeldovich mechanism
- Prompt or Fenimore mechanism
- N2O Intermediate mechanism
There are two chain reactions in the thermal or Zeldovich mechanism. A third reaction is often
included forming the Extended Zeldovich mechanism. These reactions are listed below (25):
W W W
Equation 2-21 - Zeldovich Mechanism Reaction 1
W W
Equation 2-22 - Zeldovich Mechanism Reaction 2
W W
Equation 2-23 - Extended Zeldovich Mechanism Reaction 3
Due to the large activation energy of Equation 2-21, this mechanism is strongly temperature
dependent. As a general rule, this mechanism is negligible at temperatures below 1800K (25).
38. 24
The prompt or Fenimore mechanism is closely related to the combustion chemistry of hydrocarbons.
In general, the Fenimore mechanism describes a process where molecular nitrogen and hydrocarbon
radicals react to form amines or cyano compounds. These amines or cyano compounds are
converted into intermediate compounds which eventually form NO. The Fenimore mechanism
consists of the following reactions.
W W W
Equation 2-24 - Fenimore Mechanism Reaction 1
W W W
Equation 2-25 - Fenimore Mechanism Reaction 2
In the above reaction mechanism, Equation 2-24 is the rate limiting step. For equivalence ratios less
than 1.2, the conversion of HCN to form NO follows the sequence below.
W W
Equation 2-26 - HCN Conversion Step 1
W W
Equation 2-27 - HCN Conversion Step 2
W W
Equation 2-28 - HCN Conversion Step 3
W W
Equation 2-29 - HCN Conversion Step 4
For equivalence ratios richer than 1.2, the chemistry becomes far more complex and is well beyond
the scope of this report. All mixtures tested with the catalytic burner in this study will be very lean.
Regardless of the fuel/air mixture, the fuel in this investigation is converted in a heterogeneously
catalyzed reaction and not in a gas phase reaction. This severely limits formation of the radicals
required to initiate the prompt NO mechanism.
The N2O Intermediate mechanism can be a significant source of NO in fuel lean, relatively low
temperature operating conditions. The steps of this mechanism are highlighted below (25).
39. 25
W X W X
Equation 2-30 - N2O Intermediate Mechanism Step 1
W W W
Equation 2-31 - N2O Intermediate Mechanism Step 2
W W W
Equation 2-32 - N2O Intermediate Mechanism Step 3
This mechanism is important in NO control strategies involving lean, pre-mixed combustion. These
conditions are under investigation in the field of gas turbines among others.
The previous mechanisms have been highlighted to show that the formation of oxides of nitrogen in
the catalytic burner designed for this study is very unlikely. The fuel is converted in a
heterogeneously catalyzed reaction and not in a gas phase reaction. This drastically limits the
formation of the radicals which are required in the mechanisms. Furthermore, the temperature in the
catalytic burner will never be higher than 900˚C and all three mechanisms are highly temperature
dependent. In a previous study by Wiese et al., nitric oxide emissions were measured and
determined to be negligible (3). Ditaranto et al. assumed that nitric oxide emissions were negligible
based on the findings by Wiese et al (2). As such, nitric oxide emissions will also be assumed to be
negligible for this investigation.
2.8.Unmixedness
Unmixedness was used in this investigation as a method for determining how well the fuel and air
was premixed prior to entering the catalytic core. In order to use the catalytic core most efficiently,
perfectly pre-mixed fuel and air should be evenly distributed across the entire cross-sectional area.
Once the mixture enters the core, no more mixing occurs because the ceramic monolith consists of
hundreds of thin, vertical tubes. As a result, exhaust emissions can be taken at various positions
across the core to determine local carbon concentrations for each set of experimental parameters.
Unmixedness was evaluated using the following equation from Vranos et al (26).
X
Y
Z
Equation 2-33 - Unmixedness
40. 26
In Equation 2-33, UM, σ, and μ are defined as unmixedness, the standard deviation of the total
carbon molar fraction and the mean value of the total carbon molar fraction. The equation was
simplified to allow molar concentrations from the FTIR to be directly input. The resulting equation is
shown below (Equation 2-34).
X [1
∑]^,`abcde
]afg,`abcde
[
Equation 2-34 – Simplified Local Unmixedness
In this equation, ∑Yi,Carbon refers to the sum of molar concentrations of all carbon containing
compounds at a single point in the exhaust stream. In this investigation, an exhaust probe was
moved horizontally across three fixed positions (center, quarter and outer) for every set of
experimental parameters to obtain an emissions profile across the core. Yave,Carbon refers to the
average molar concentration of carbon for each set of experimental parameters. The average value
was calculated by adding the 3 summations of molar carbon concentrations (from the center,
quarter and outer positions) and then dividing the total by 3.
A low value of unmixedness indicates that the fluids were well mixed across the reactor. A high value
of unmixedness implies that large gradients in species concentration exist. In this investigation, the
equivalence ratio for each fuel was kept constant.
2.9.Previous Studies
While catalytic combustion has been investigated at length, only a small number of studies have
directly looked at the catalytic combustion of fuel cell anode waste gas. Catalytic combustion is a
complex area of study involving thermodynamics, mass and heat transport as well as chemical
kinetics (27). The complexity is further increased when multi-component fuels are utilized, as the
different components can either promote or reduce catalytic activity due to increased competition
for surface sites (28). Thorough premixing of the fuel and oxidizer is also critical for these systems, as
catalytic combustion is a surface reaction process and poor mixing can increase emissions due to
inefficient use of the available reaction sites (3).
Numerous studies were analyzed but this investigation shares the most similarities with studies by
Ditaranto et al. and Wiese et al. (2) (3). Both studies involved the analysis of the exhaust products of
catalytic combustion of anode waste gas. The study by Ditaranto et al. was particularly relevant to
41. 27
this investigation as this group also investigated the exhaust products of the catalytic combustion of
evaporated methanol.
Both studies concluded that hydrogen is readily consumed on the surface of the catalytic core and
that the main challenges reside in burning off the minor species of the anode waste gas. The
predominant minor species included methanol slip and carbon monoxide from the steam reforming
process. Both studies assumed an axi-symmetric reactor and this assumption will be continued here.
Wiese et al. measured NOx emissions and determined them to be negligible for many of the
aforementioned reasons described in detail in Section 2.7. Ditaranto et al. neglected measuring NOx
emissions citing Wiese et al. and another study by Kuper et al. discussing a catalytic combustor
concept for gas turbine engines (29).
The catalytic core in the experimental setup by Wiese et al. featured a hollow ceramic cylinder
wrapped in a ceramic-coated wire mesh as shown in Figure 8. In this concept, the premixed fuel and
air pass through the ceramic layer and then react on the catalyst-coated wire mesh. The back
pressure caused by the porous ceramic helped create a uniform flow across the mesh. The wash coat
on the metal mesh consisted of γ-Al2O3 with a platinum catalyst. As in this investigation, a maximum
operating temperature of 900˚C was adhered to by Wiese et al. In their investigation it was noted
that exceeding this temperature caused a phase change from γ-Al2O3 to α-Al2O3 decreasing the
surface area by a factor of 60 (3).
42. 28
Figure 8- Wiese et al. Experimental Setup (3)
Wiese et al. provided a fundamental, yet significant, insight into the emissions related to catalytic
combustion of anode waste gas. There were some notable differences however between their work
and the investigation outlined in this report.
1. The burner design in this investigation is far simpler, as this is necessary for a production-type
system. Furthermore, it has far less backpressure than the burner used by Wiese et al. as it is
a flow through design with larger channels.
2. The air and fuel are not preheated as they were in Wiese et al. In a commercial system, it
would be very difficult to preheat the incoming air to 200˚C, especially since most of the
excess heat is used for other purposes.
3. A unique, proprietary catalyst composition was used in this investigation. A comparison in
composition cannot be made with Wiese et al. on this parameter as the catalyst is treated as
a black box in this investigation.
4. The fuels used in Wiese et al. were either pure methanol or various mixtures of methanol and
hydrogen. Direct comparisons can only be made for the methanol tests in this investigation.
The study by Ditaranto et al. featured an experimental setup which shared more similarities with the
setup in this investigation. The setup is shown below in Figure 9. It featured a similar circular plug
style flow through monolith with tiny channels running the length of the reactor, a pre-mixer for the
fuel and air in front of the reactor and well insulated walls surrounding the catalytic core. Although
43. 29
not explicitly stated in the report, it appears that the catalyst core used was limited to a maximum
temperature of approximately 800˚C. As their tests did not investigate the emissions from pure
catalytic combustion of methanol, comparisons cannot be made.
Figure 9- Ditaranto et al. Experimental Setup
The experimental apparatus used by Ditaranto et al. was far closer in design to the one in this
investigation however some significant differences still remain. They are listed below (2).
1. Ditaranto et al. used an inline mixer to combine the fuel and air. This investigation uses an
inline mixer as well except mixing is used as an experimental variable. Ditaranto et al. makes
the assumption that the fuel and air are well mixed prior to entering the catalytic reactor.
2. The air and fuel are preheated as in Wiese et al. As mentioned above, this is not practical for
an experimental setup trying to emulate realistic field conditions.
3. While the shape of the catalytic monolith is similar to the one used in this investigation, it is
much smaller in scale (50% smaller radius, 50% shorter in length) and the catalytic
composition cannot be directly compared, other than to simply state that it is different.
4. Catalytic combustion of pure methanol was not investigated by Ditaranto et al.
44. 30
5. The composition of the simulated anode waste gas used by Ditaranto et al. featured a
significantly greater hydrogen content and a lower CO content.
Ditaranto et al. argued that was necessary to use a greater percentage of hydrogen in their
simulated anode waste gas mixture because a lower hydrogen content would have reduced the
reactor temperature. However, the reactor temperature can be directly adjusted by altering the
amount of air in the mixture. In this investigation, a simulated anode waste gas mixture was chosen
based on a worst case scenario. It features roughly twice the amount of CO, approximately 20% less
hydrogen and will be detailed in a later section.
Both Wiese et al. and Ditaranto et al. assumed an adiabatic reactor with axisymmetric properties.
These assumptions will also be used in this investigation. A review of the literature failed to find a
study that directly addressed the unique requirements of Serenergy:
1. A new, proprietary catalyst composition
2. Non-preheated air
3. A unique anode waste gas composition which includes a relatively high percentage of CO
4. A larger scale of burner
5. A unique burner design with low backpressure and modular components
These requirements, all of which are critical to the design of a proprietary system within the
company, validate the necessity for an independent investigation.
45. 31
3. Design
Mobile HTPEM/SR power packs running on methanol, such as the system depicted in Figure 4, are
one of many possible emission reduction solutions available to modern automotive manufacturers.
Due to the low reforming temperature of methanol with current Cu/ZnO/Al2O3 catalysts and the high
H/C ratio, it is an ideal fuel for use in these systems. While several manufacturers, including Daimler-
Chrysler and General Motors, have attempted to commercialize methanol reformer/fuel cell cars in
the past, none had the luxury of using an HTPEM fuel cell with PBI membranes. As a result they
required complex reformate gas cleanup systems. These systems greatly increased the complexity
and reduced the efficiency of an otherwise commercially viable system.
Steam reforming of methanol is relatively well understood, in part due to the exceptional work of
Peppley et al. (10) (11). HTPEM fuel cell systems are also reasonably well understood with Serenergy
A/S of Denmark generating a significant market presence since their inception in 2003. System
integration is an area of study where a lot of work remains however HTPEM/SR systems are currently
being used in combination with DC/DC converters connected to battery packs with positive results.
When using reformate gas (ie. not pure hydrogen) in any PEM fuel cell, the anode side of the stack
must remain open, otherwise the non-hydrogen gases will build up in the system and potentially
starve the membranes of fuel. Also, it is nearly impossible to use all of the hydrogen in the reformate
gas stream when the fuel cells are stacked because the gas generally has to pass through a channel
system on the anode side of the membrane. As the anode gas passes through the channel laterally
across the membrane, some of it is absorbed through the GDL and into the membrane. When the gas
reaches the end of the channel, there is a much lower percentage of hydrogen remaining in the
stream. The percentage of remaining hydrogen can never be allowed to reach zero as membrane
starvation would be the inevitable result. For the system under consideration in this study, the
percentage of hydrogen in the reformate stream is approximately 30%. This number is an optimized
value based on the thermodynamic needs of the entire system and will be taken as a fixed value for
the purposes of this study.
A key component in generating a highly efficient HTPEM/SR power pack is the burner as this is the
component which converts the leftover hydrogen, methanol slip and CO in the anode waste gas
(AWG) into useable heat to support the endothermic methanol-steam reforming reaction. This
component faces some unique demands as the equivalence ratio of the fuel/air mixture can vary
46. 32
widely and rapidly depending on the demands of the system. A design using catalytic combustion
was chosen as opposed to using a non-premixed diffusion flame because heat release can be
sustained at low temperatures over a wide range of equivalence ratios while maintaining low
emissions (2). The variability of the AWG stream both in terms of composition and volume would
make it difficult to maintain homogeneous combustion let alone optimizing the combustion for low
emissions. Also, a catalytic burner can take many forms allowing for flexibility with respect to system
integration.
Another design requirement which supports the use of a catalytic burner is the ability to combust
evaporated methanol when the system is cold in order to preheat both the fuel cell stack and the
reformer. The method for distributing the heat is outside the scope of this report but requires an
exhaust temperature of less than 600˚C. The same catalytic core will be used for the combustion of
both fuels with no modifications.
While the control system design is outside the scope of this study, a range of equivalence ratios were
considered in the design of the burner. In a real system, there are fluctuations in the amount of
hydrogen consumed in the anode which cause unavoidable temperature peaks or valleys due to the
lag time in air blower response. To reduce complexity, these variations were not emulated however
the maximum expected fuel flow rate was used throughout testing to simulate a worst case type
scenario.
3.1. Design Parameters
In order to test the emissions products of catalytic combustion, an experimental setup was designed
which allowed for the measurement of many different variables simultaneously. The setup was
designed to mimic the conditions the combustor would see if used in a commercial combined
HTPEM/methanol-steam reformer system. These conditions include:
- Air pressure close to atmospheric
- Anode waste gas simulation (see Table 3 for composition) pressure slightly above
atmospheric
- Fully evaporated methanol at slightly above atmospheric pressure
- Startup from room temperature using pure methanol
47. 33
The finished design will ultimately be used to test the emissions from the Johnson Matthey HiFUEL-
AB4 catalyst using two different fuels. The first fuel tested will be pure, evaporated methanol. The
second fuel tested will be an anode waste gas simulation with the composition listed in Table 3. The
oxygen supply in both cases will consist of non-preheated air.
Component Ratio
H2 0.3158
CO2 0.6568
CO 0.0274
Table 3 - Anode Waste Gas Composition
Real anode waste gas in Serenergy’s HTPEM/SR system actually contains water vapour however a
recent study by Wiese et al. stated that carbon dioxide was an acceptable substitute as the water did
not actively participate in the reaction (3). Wiese et al. also noted that excess carbon dioxide in place
of the water tended to increase carbon monoxide emissions slightly. Replacing the water vapour
with carbon dioxide should produce worst case emissions results.
The goal was to provide a premixed fuel/air mixture over the entire cross section of the catalyst
while avoiding any open flame because once the mixture hits the catalyst, there is no more mixing.
An open flame would raise the temperature of the catalyst above the 900˚C maximum
recommended by the manufacturer. The catalyst is coated onto a ceramic monolith with narrow
channels running the length of the core (see Figure 10 below). These channels are what prevent
further mixing. For the purposes of this investigation, the catalyst core was treated as a ‘black box’.
Figure 10 - Catalytic Monolith
The entire system contained enough internal adjustability to adapt to the mixing requirements of
either fuel. It was also modular so that it was relatively easy for the end user to swap out the catalyst
core in order to test different catalyst compositions. Modularity would also allow the end user to
48. 34
make adjustments to the final design as more is learned about the system. Some final design
requirements included a fully automated fuel evaporator and data collection system as well as active
and passive safety features. The entire test bench was also portable so that it could be delivered to
Serenergy in Denmark upon completion of the project for testing of future catalyst compositions.
3.2. Moveable Block Swirl Generator Design
A moveable block swirl generator was chosen to premix the fuel and air as it can provide a wide
range of adjustability with respect to turbulence. Also, this type of swirl generator has the ability to
form greater swirl intensity with lower pressure drop than other types of swirl generators (21) (20).
Pressure drop is an important consideration as moving air through a system is energy intensive. The
lower the pressure drop, the higher the overall system efficiency.
Too much swirl could cause undesired back pressure to build up in the bulk flow due to the
formation of a central recirculation zone (20). If a central recirculation zone forms when the
simulated anode waste gas is used, there is potential for non-catalytic combustion in the mixing
chamber. This must be avoided as the purpose of the swirling flow is for mixing, not for combustion.
The critical point at which the low pressure region in an isothermal (cold) flow is able to form a
central recirculation zone is generally between 0.5 ≤ S ≤ 0.6 (20) (21). This value refers to the actual
swirl value present in the burner. Equation 2-19 has been developed assuming a uniform axial
velocity distribution at the plane of interest highlighted in Figure 6 (21). Also, since the derivation
does not include the contribution of the radial pressure distribution in the fluid on the axial
momentum flux (Equation 2-16), S is not conserved along the length of the mixer/burner (30) (22). In
a confined chamber, such as the one used in this design, there is a loss of angular and axial
momentum due to friction effects with the walls (30). In this case, there is an even greater loss in
axial momentum due to the flow restriction caused by the catalytic monolith. Due to the above
mentioned effects, it is likely that the actual swirl number in the region of the injector jets is
significantly lower than the theoretical value calculated at the exit plane of the swirl blocks. As a
result, the swirl numbers used in this report should only be used as a measure of relative swirl
intensity.
With the knowledge that swirl intensity diminishes along the length of the mixing chamber, a
moveable block swirl generator was designed which can create flows with 0 ≤ S ≤ 4.85. In order to
49. 35
generate this range of S, the values listed in Table 4 were chosen for the moveable block swirl
generator design specifications. These values correspond with the geometry shown in Figure 7.
Parameter Description Value
n Number of swirl blocks 3
R Swirl generator exit radius 25.4mm
Rh Swirl generator inner radius 3.97mm
B Depth of swirl blocks 31.75mm
α Fixed swirl block angle 55 degrees
ξ Adjustable swirl block angle -
ξm Maximum opening angle 12 degrees
Table 4 - Moveable Block Swirl Generator Design Specifications
The final design was machined out of 6061 aluminum and is shown in the maximum S position below
in Figure 11. It is installed on the bottom of a mixing chamber which is highlighted in a later section of
the report.
Figure 11 - Completed Moveable Block Swirl Generator
3.3. Momentum Flux Ratio
The momentum flux ratio J was a very important consideration in the design of the mixing chamber.
The theory behind mixing of non-reactive gaseous flows was examined in the literature review
section previously. As is the case in many real world applications, the design of the mixing chamber is
not a perfect representation of the theoretical case. The equations in the literature examine two
distinct cases (13):
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1. Single sided injection in a rectangular cross-flow (Equation 2-11)
2. A circular duct with circumferential injectors pointing inwards (Equation 2-15)
The burner designed for this investigation features a cylindrical fuel injector with the injectors
pointing outwards into a cylindrical duct. See Figure 12 below for a representation of the three
distinct cases. Due to the relatively low flow rates involved and the packaging requirements of the
burner, it would have been very difficult to design a burner with circumferential jets pointing
inwards. With the chosen design, it was also very easy to change the jet diameter or the number of
jets without disassembling the entire setup. As this is primarily a test rig for Serenergy to test various
fuels and catalysts, this was a very important consideration.
Figure 12 - 3 Distinct Injector Patterns
From the three cases listed above, Jopt for single sided injection into a rectangular cross-flow is
greater than Jopt for circumferential injection into a cylindrical cross-flow for the parameters used in
this investigation. These equations are merely guidelines. As no data was available specifically
relating to the unique design (Actual Case) on the right, and no correlations were available in
literature, an estimation was made on the calculation of Jopt. If the ‘Actual Case’ shown in Figure 12
was unrolled, it would look something like single sided injection into a rectangular duct case.
Assuming the same number of jets in the calculation of Jopt in Equation 2-11 and Equation 2-15, Jopt for
single sided injection into a rectangular cross-flow is greater than Jopt for circumferential jets pointing
inwards into a cylindrical cross-flow over the range of conditions examined. Since the actual case is a
mixture of the two theoretical cases, it was decided that Jopt for the mixing chamber should be
somewhere in between the two assuming the same number of jets.
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To further complicate matters, swirl is present in the cross-flow. In general, when swirl is present in
the cross-flow, Jopt is higher than in the equivalent, non-swirl case (16). The multiplicity of parameters
makes it difficult to derive appropriate correlations so it was determined that Jopt should sit at
roughly the mid-point of the two theoretical cases. The calculation process for Jopt using methanol as
a fuel is detailed in Table 5 below. J_Experimental was calculated using Equation 2-9 from a previous
section. An identical spreadsheet was created for the simulated anode waste gas. Two separate
injectors were ultimately machined, one for each fuel. Table 5 below shows the values for methanol.
Methanol
Q (Gaseous Fuel) 7.50 LPM
Hole ø 0.055 in
Hole ø 0.001397 m
Hole Area 6.131E-06 m^2
n (# of Holes) 6
Total Area 3.679E-05 m^2
V_jet 3.398 m/s
rho_jet 1.007 g/L
V_jet 3.398 m/s
rho_main 1.184 g/L
V_main 0.695 m/s
J_Experimental 20.308
J_opt (Circular) 11.520
J_opt (Rectangular) 22.797
Table 5 - JOpt Calculations
When the fuel injector jets on the methanol injector were machined, they ended up slightly oversize.
The intention was to create uniform hole diameters of 0.055” however the final hole diameters were
closer to 0.060”. This resulted in a J_Experimental value of 14.33 which is between the two
theoretical cases. Since this was an estimated parameter, it was decided to use the injector as
machined and record the results. Future injector jets will be machined using electrical discharge
machining. This will allow a much more accurate hole size to be created. Final J_Experimental values
for both fuels are listed below in Table 6.
Fuel J_Experimental Value
Methanol 14.33
Simulated AWG 83
Table 6 - J_Experimental Values
The injector for the simulated AWG testing was machined after the conclusion of the methanol
testing. The J_Experimental value for the simulated anode waste gas test was set at 83 for reasons
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that will be discussed in the Results and Discussion section. Due to the number of holes required to
keep the momentum flux ratio at a reasonable value and the small diameter of the fuel inlet rod, it
was necessary to stagger the holes slightly in this case.
3.4.System Components
The experimental apparatus featured a modular design. This way certain sections could be swapped
out to meet specific test requirements. The components which required custom designing are listed
and described in detail below. All custom designed machined component drawings are shown in
Appendix A. All purchased components are listed in Appendix B. A diagram of the catalytic burner
assembly is shown in Figure 13 for reference. A schematic of the entire lab bench setup is shown in
Figure 18.
Figure 13 - Catalytic Burner Assembly
3.4.1.Evaporator
An evaporator was designed to vaporize up to 0.6L/ hour of pure methanol fed from a peristaltic
pump. This is in excess of the expected maximum flow rate of 0.45L/hour. It was machined from
6061 aluminum and anodized for corrosion resistance. It was controlled using a PID temperature
53. 39
feedback loop through LabVIEW combined with a relay to turn the heating cartridges on and off as
required to maintain a specified set point regardless of fuel flow rate. The LabVIEW 0V (off) or 5V
(on) signal was sent through a National Instruments USB 6009 multifunction DAQ. From there it
triggered a standard electromechanical relay which turned on the 24V DC heating cartridge power
supply. The 100W cylindrical heating cartridges were placed above and below the channels inside the
evaporator for evenly distributed heating. The entire assembly was wrapped in high temperature
insulation. Prior to being wrapped in insulation, the unassembled evaporator can be seen in Figure 14
below. Upon leaving the evaporator, the vaporized fuel passed through a heated stainless steel tube
and into the mixing chamber.
Figure 14 - Finished Evaporator Prior to Assembly
3.4.2. Mixing Chamber
This chamber was designed using guidance from previously referenced papers on swirl generators
and the mixing of jets in confined subsonic cross flows. The purpose of this chamber is to provide a
well premixed mixture of fuel and air for the catalytic monolith module below. The fuel injector
height can be altered to change the amount of mixing time available and the moveable block swirl
generator can be adjusted during operation to alter the value of S. As two different fuels were used,
this adjustability reduced the amount of setup time between tests. See Figure 15 below for a
representation of the mixing chamber.
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Figure 15 - Mixing Chamber
There are three, equally spaced air inlets on the bottom of the outer air guide. Due to the lack of
pressure drop across the swirl generator, it was decided that a single air inlet could potentially
favour one of the three windows of the swirl generator. Also, the bottom half of the swirl generator
is bolted to the outer air guide. Swirl is adjusted by twisting the outer air guide.
3.4.3. Catalytic Monolith Module
The catalyst used for this investigation was a proprietary commercial blend from Johnson Matthey in
Reading, United Kingdom coated onto a ceramic monolith. The mixture is platinum based and
designed specifically for Serenergy’s unique requirements (pure methanol at first, switching to
anode waste gas after several minutes of operation). Several identical monoliths were ordered to
validate the emissions results. The catalytic monolith was wrapped in 3M Interam and placed inside a
304 stainless steel burner tube. The Interam expands permanently the first time it is heated to block
gases from leaking around the edge of the monolith. The finished catalytic monolith module is
shown below in Figure 16. You can see the layer of 3M Interam between the ceramic monolith and
the steel wall of the burner tube.
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Figure 16 - Catalytic Monolith Module
3.4.4. Catalyst Core
The catalyst core used for this investigation was a coated ceramic monolith with roughly 0.125”
square channels running the entire length. Figure 16 shows these channels clearly. The wash coating
consists of platinum, rhodium and palladium in less than 1 weight percent quantities as well as a
roughly 10 weight percent mixed metal oxide component. It was sourced from Johnson Matthey and
is classified as their HiFUEL – AB4 mixture. The exact composition was not relevant for this
investigation as the purpose was to test the catalyst core as a finished product to evaluate emissions
characteristics. The two constraints the manufacturer had on using the coated monolith were to
keep the operating temperature below 900˚C and the space velocity below 50000hr-1
. The maximum
temperature seen in actual use will be approximately 600˚C so the only requirement which needed
careful consideration was the space velocity. The space velocity is calculated as follows.
8 9 # 9hi
#j k / 9 1' 8 6 j6
#j k X1hl
Equation 3-1 - Space Velocity
An iterative approach to overall system size was completed where a maximum space velocity of
50000hr-1
was a key constraint. The maximum fuel and air flow rates possible using the available
mass flow controllers also played a role. After a considerable number of iterations, a catalyst core
size of 2” in diameter and 5” in length was chosen. See the Discussion section for an explanation of
the final decisions. At the maximum space velocity, the core should be able to efficiently burn
56. 42
0.45L/hour of methanol or 26L/min of simulated anode waste gas. Equilibrium temperatures for the
design study were based on results generated by the STANJAN code (31).
3.4.5. LabVIEW Visual Interface
All of the components were calibrated and controlled through a custom LabVIEW visual interface.
Version 8.6 was used for this investigation. Data for all input and output channels were automatically
logged in Microsoft Excel for every test. A sample front panel view of the visual interface is shown
below in Figure 17.
Figure 17 - LabVIEW Visual Interface Front Panel
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4. Experimental Methodology
This section provides a detailed description of the experimental setup used to study the emissions
characteristics of the two fuels under investigation. This is a unique setup designed with modularity
and ease of adjustment in mind.
4.1. Final Experimental Setup
A schematic of the finished system is shown in Figure 18 to highlight the basic overall layout of the
experimental setup.
Figure 18 - Experimental Setup
