ARTICLE IN PRESS
Chemical Engineering Science 65 (2010) 3709–3717
Contents lists available at ScienceDirect
Chemical Engineering Science
journal homepage: www.elsevier.com/locate/ces
Manganese oxide dissociation kinetics for the Mn2O3 thermochemical
water-splitting cycle. Part 1: Experimental
Todd M. Francis, Paul R. Lichty, Alan W. Weimer Department of Chemical and Biological Engineering, University of Colorado, 1111 Engineering Dr., Campus Box 424, Boulder, CO 80309, USA
a r t i c l e in fo
abstract
Article history:
Received 3 July 2009
Received in revised form
20 February 2010
Accepted 2 March 2010
Available online 6 March 2010
It is shown that the dissociation of Mn2O3 to MnO in a short residence time aerosol flow reactor can
achieve high conversions approaching 75% when the concentration of oxygen is kept below 0.25%.
Significant recombination reaction occurs when the oxygen content exceeds 0.25% by volume. A dual
reaction mechanism for Mn2O3 dissociation was found:
Keywords:
Chemical reactors
Reaction engineering
Kinetics
Energy
Manganese oxide
Thermochemical water splitting
RAvramiErofeev ¼ A1 eEa,1 =RT nð1XÞ½lnð1XÞðn1Þ=n
ROrder_of _reaction ¼ A2 eEa,2 =RT ð1XÞn
with the transition from one mechanism to the other occurring at an extent of reaction of
approximately 0.6. Rate constants for the two mechanisms were calculated to be 1.8 107 7 1.3 107
and 5.6 103 7 4.1 103 s 1, respectively, for oxygen concentration o 0.25%. High levels of dissociation
are achievable when the reaction is carried out in an inert gas environment using a reactor
configuration that limits the reverse reaction.
& 2010 Elsevier Ltd. All rights reserved.
1. Introduction
Renewable energy is a reliable method to produce energy that
is sustainable and environmentally friendly. In the past, energy
from renewable sources has not been in the forefront of
discussions. However, with recent increases in the price of fossil
fuels, renewable energy is becoming an increasingly popular
topic. Although the technology is still developing and not yet
economically practical, many believe that the future of renewable
energy is hydrogen.
A three-step thermochemical water-splitting cycle to generate
renewable hydrogen has been proposed, which uses manganese
oxide and solar-based thermal energy to produce hydrogen
(Nuesch, 1998; Sturzenegger et al., 1999; Sturzenegger and
Nuesch, 1999):
Mn2O3
solar energy
!
2MnO+ 1/2O2
(1)
2MnO+ 2NaOH-H2 + 2NaMnO2
(2)
2MnO2 + H2O-Mn2O3 + 2NaOH
(3)
H2O-H2 + 1/2O2
(4)
Corresponding author.
E-mail address: Alan.Weimer@Colorado.edu (A.W. Weimer).
0009-2509/$ - see front matter & 2010 Elsevier Ltd. All rights reserved.
doi:10.1016/j.ces.2010.03.002
An efficiency study of this cycle was performed, where
efficiency was defined as the amount of energy needed to split
water into its constituents over the amount of energy input to the
entire cycle (Sturzenegger and Nuesch, 1999). The maximum
efficiency, which was the efficiency of an idealized system, was
calculated to be 74%. When all unit operations and heat recoveries
were included, the efficiency was then between 16 and 22%. This
study lends optimism that this cycle has the potential to supply
sustainable hydrogen efficiently.
Most work on the sodium manganese oxide cycle has focused
on the hydrogen generating step (Eq. (2)) and the product
recovery step (Eq. (3)) with little work done on the reduction
step (Eq. (1)) (Nuesch, 1998; Sturzenegger et al., 1999). It is
essential to understand the reduction step because the entire
feasibility of this thermochemical cycle is based on this reaction
having a high overall conversion (Sturzenegger and Nuesch,
1999). The viability of preventing MnO and O2 recombination
must be investigated in order to determine the degree to which
recombination impacts the overall process (Ganz et al., 1998;
Sturzenegger et al., 1999; Perkins and Weimer, 2004). A
high-temperature aerosol flow reactor (AFR) was used to study
the thermal dissociation of manganese oxide. This type of dilute
phase particle phase reactor has shown to yield high dissociation
reaction rates using other metal oxide systems. With the inherent
importance of the thermal dissociation step (Eq. (1)), a rapid
dissociation and maximum conversion are essential for the
viability of a high-throughput thermochemical cycle for efficient
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T.M. Francis et al. / Chemical Engineering Science 65 (2010) 3709–3717
hydrogen production. Additionally, an AFR can easily be outfitted
with a low-temperature quench zone, which can be used to
rapidly cool the particles upon leaving the high-temperature
zone. This quench zone reduces the potential for recombination
and helps the reaction conditions to be tightly controlled.
The dissociation of manganese oxide is reversible, as are
most high temperature reduction steps from metal oxide
thermochemical cycles (Otto, 1964; Matsushi and Thoburn,
1965; Ganz et al., 1998). In a proof of concept study, it was found
that when Mn3O4 is reduced, MnO is only formed if the reaction
was quickly quenched (Ganz et al., 1998; Sturzenegger et al.,
1999). Conversions were between 5 and 15% MnO when the
reaction was quickly quenched and 0% MnO when the reaction
was not. Sturzenegger et al. were unsure if the reaction conditions
or the reactor design led to low conversions (Sturzenegger et al.,
1999), however a fast quench in a thermal gravimetric analyzer
(TGA) was shown to increase the conversion of Mn2O3 to Mn3O4
from 34% to 95% (Otto, 1964). This result lends optimism that if a
reactor can be designed to give a fast enough quench, high
conversions may be retained. It was for this reason that an AFR
was hypothesized to be a reactor capable of achieving high
conversions.
There are only a few examples of high-temperature AFRs for
the processing of solids found in the literature, but those
examples do provide insight into why an AFR would work well
for the dissociation of manganese oxide. An AFR has been used to
rapidly produce submicron sized boron carbide particles directly
at a high operating temperature (Weimer et al., 1991). Particles
that were uniform in size were produced at 2200 K via a
rapid quench that limited grain growth. Residence times were
between fractions of a second and only a few seconds. An
expanded cooling zone, or quench zone, was used to effectively
stop the reaction. Rapid cooling is a method that may help reduce
the potential for recombination during the dissociation of
manganese oxide.
An AFR has also been used for reacting aluminum particles with
nitrogen gas at 1873 K to form aluminum nitride (Weimer et al.,
1994). The reaction was carried out to high conversions with little
deposition of particles on the walls. This indicated that heating rates
and residence times attainable in an AFR were sufficient for
gas–solids reactions to be driven to high conversions.
Carney et al. used an AFR for the decomposition of nickel
oxalate (Carney et al. 2006a, b). Nickel oxalate was fed into the
reactor and reduced in the presence of hydrogen to produce nickel
powder. High conversions were achieved within short residence
times, and were largely attributed to the fast heating rate in the
reactor. Additionally, the product particles consisted of nanosized primary particles, which are desirable in thermochemical
water-splitting cycles. This shows an additional advantage of
studying the dissociation in an AFR.
A solid particle receiver has been demonstrated to efficiently
heat solids using concentrated solar energy (Hruby et al., 1984). A
model of a flowing cloud of particles showed that temperatures of
up to 1118 K could be achieved for micron size particles (Evans
et al., 1987). Further, Dahl et al. showed an ‘‘on-sun’’ solar reactor
achieving temperatures greater than 2000 K (Dahl et al., 2004a, b).
Methane gas was dissociated to form carbon black and hydrogen
and conversions up to 90% were achieved for a residence time of
0.1 s (Dahl et al., 2004a, b).
Finally, an AFR has been used to study the reduction of metal
oxides as well. ZnO was split into its constituent parts in an
AFR at high temperatures (1873–2073 K) and over short residence
times (1–2 s) (Perkins et al., 2008). High conversion to Zn
metal was achieved, and the resulting primary particles were
precipitated out of the gas phase (Zn was a vapor) and were
nano-sized.
These studies showed that an AFR has many advantages for
dissociation reactions including rapid heating rates and the ability
to control short residence times followed by rapid cooling to
prevent particle growth. Hence, the AFR configuration provides
for an opportunity to dissociate manganese oxide and to
potentially utilize rapid cooling to limit the reverse reaction from
occurring. Further, these types of reactors have been shown to be
scalable for industrial production.
2. Materials and methods
2.1. Equipment
Kinetic studies were completed using a Thermal Technology
Inc. AFR, Fig. 1. The hot zone was maintained using a high-density
graphite resistance heating element combined with high-purity
graphite felt insulation that reduced heat losses. The reactor tube
placed in the center of the furnace was a 99.8% slip cast alumina
protection tube (ID¼0.09 m, L¼1.17 m) from McDanel Advanced
Ceramic Technologies LLC.
The alumina tube AFR has the capability to achieve reaction
temperatures as high as 2173 K, limited by the upper temperature
for using alumina. A Eurotherm 914 controller was utilized for
temperature control within the reactor. Temperatures were measured on the wall of the hot zone using a Type-C thermocouple up to
Sweep
Nitrogen In
Fluid Bed Feeder
Feed Tube
Hot Zone
Graphite
Element
Quench Tube
Alumina Tube
Outer Cooling Zone
HEPA Filter
To Gas Analyzers
and Vent
Gravity Collection Vessel
Fig. 1. Schematic of the aerosol flow reactor.
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1073 K. From 1073 to 1373 K, the temperature measurement from
the thermocouple was blended with that from an infrared
pyrometer (Ircon Inc.) for seamless temperature control over the
intermediate range. Above 1373 K, the temperature was measured
only with the infrared pyrometer.
All powder that was fed to the reactor and processed was then
quickly quenched in a quench tube assembly, shown in Fig. 2. This
quench effectively ‘‘terminated’’ the reaction for a more accurate
calculation of the time the particles resided in the hot zone. The
quench process also effectively inhibited the recombination
reaction. The quench tube consisted of concentrically-mounted
inner and outer copper tubes. The reason for this was to ensure
uniform cooling along the copper tube. Because the quench tube
was partially inserted into the hot zone, a siphon/pump system
was employed to ensure that cooling water was always flowing.
The copper tube was wrapped with zirconia insulation for
thermal protection; a zirconia cap was mounted on the top of
the quench assembly for further insulation.
After the powder passed through the quench tube, it was
collected in a gravity collection vessel. Any powder that remained
entrained and left the collection vessel was trapped by a high
efficiency particulate air (HEPA) filter. The filter was manufactured by General Electric H2O Process Technologies, had a 0.22 mm
pore size, and was 0.14 m in diameter. After passing through the
filter, gas was analyzed in an oxygen analyzer (Advanced Micro
Instruments Model 201) and was subsequently vented. The
Zirconia Cap
Brass Weir
Tube
Inner Copper
Tube
Outer Copper
Tube
Zirconia Cloth
Insulation
Outer Cooling
Zone
3711
oxygen analyzer had the ability to measure oxygen concentration
over seven ranges, from 1000 ppm to 100%. A fluidized bed feeder
was used to feed dispersed particles into the reactor.
2.2. Experimental plan to investigate reaction without oxygen
The first experimental plan to study the kinetics of Mn2O3
dissociation was performed in an inert environment. This was
done to determine if high conversion to MnO was feasible in the
limited residence times in the reactor. Additionally, there are
available rate laws for the dissociation without oxygen (Francis,
2008). The dissociation to MnO occurs in two reactions where
Mn3O4 is an intermediary (Eqs. (5) and (6)).
3Mn2O3-2Mn3O4 + 1/2O2
(5)
2Mn3O4-6MnO+ O2
(6)
Mn2O3 powder (Sigma-Aldrich, 99% purity), having a primary
particle size of less than 44 mm was used to investigate the
dissociation. Temperature and bulk gas flow rate were used as
factors to study the kinetics of the dissociation. A central
composite design (CCD) with the two factors (temperature and
bulk gas flow rate) was performed where temperature was varied
between 1673 and 1873 K and the bulk gas flow rate between 2.5
and 7.0 slpm. The lower temperature was selected because
FACT thermodynamics software (Facility for the Analysis
of Chemical Thermodynamics; McGill University) predicted
significant conversion; the higher temperature was selected
because it was approximately at the lower limit of the dissociation of popular two-step thermochemical cycles. If the reaction
temperature for the dissociation of Mn2O3 needed to be much
higher, then an alternative two-step water splitting cycle would
likely become more desirable, due to inherent simplicity and
potentially better process efficiency. The gas flow rate range was
selected because initial calculations indicated that it would
provide a desirable range of residence times ( 1–3 s).
At the completion of an experimental run, the product powder
was analyzed for oxygen content with a LECO Corporation TC600
oxygen determinator. X-ray diffraction (XRD) (Cu-k(a)) was used
to identify product crystallinity, which helped to identify which
forms of manganese oxide were in the product. By combining the
results from the TC600 oxygen determinator and XRD, product
conversion was calculated. See Appendix A for the calculation
steps. In addition, a titration technique was used to determine the
Mn + 2/Mn + 3 ratio. The titration technique supported conversions
calculated from oxygen percent.
2.3. Experimental plan to study oxygen recombination effect
An inherent challenge with the dissociation of metal oxides is
the potential for recombination to occur in the quench step. In
order to better understand the recombination effect in an AFR,
two different experimental plans were executed. The first was a
study of the recombination of MnO with O2. Eqs. (7) and (8) show
the expected recombination reactions.
Hot Water
Out
Cold
Water In
To Collection
Vessel
Fig. 2. Schematic of the quench tube assembly.
6MnO+ O2-2Mn3O4
(7)
2Mn3O4 + 1/2O2-3Mn2O3
(8)
A 200 mesh, 99% MnO powder from Alfa Aesar was used to study
the recombination. The same factors from the experimental plans
in Section 2.2 (temperature, gas flow rate) along with the addition
of oxygen, were used. A three-factor CCD was employed over
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temperatures of 973–1573 K, bulk gas flow rates between 2.5 and
7.0 slpm and oxygen concentration between 0.25 and 1.0%. The
temperature range was selected along with the oxygen concentration because these range combinations were predicted by FACT
to give significant MnO oxidation at safe conditions in the
reactor. Again, oxygen percents from the LECO TC600 oxygen
determinator were used to calculate conversions.
To add to the understanding of the recombination of MnO, an
experimental plan of the dissociation of Mn2O3 was performed,
with the addition of oxygen to the bulk gas flow. A 23 full-factorial
design was executed with temperature, gas flow rate, and oxygen
content as the factors. The ranges of temperature and gas flow
rates were the same as those used in the experimental plan
described in Section 2.2. The range for oxygen percentages
(0.25–1.0%) was identical to that used during recombination
studies.
3. Results and discussion
3.1. Mn2O3 to MnO forward reaction kinetics
A typical oxygen generation curve of the dissociation reaction
is presented in Fig. 3. The reactor temperature was 1873 K and
had a bulk gas flow rate of 7.0 slpm. Several peaks may be seen in
the center of the graph. These resulted from a small amount of
vibration being applied to the feeder, which periodically increased
the amount of powder fed to the reactor. Additionally, there were
two discontinuities in the graph. The discontinuities occurred
when the change in oxygen was so rapid that the O2 sensor range
was not changed quickly enough. It is evident from Fig. 3 that the
amount of oxygen in the system rapidly increased when
manganese oxide was fed.
All of the Mn2O3 powder from the experimental runs reacted
completely through the first dissociation (Eq. (5)) and partially
through the second dissociation (Eq. (6)). XRD data shown in Fig. 4
verified this, as the product spectrum showed the presence of
both Mn3O4 and MnO spectra, with no evidence of peaks from the
Mn2O3 spectrum. From the derivation in Appendix A, the
conversion to MnO as a function of O% present in the sample is
given as
XMnO ¼
2MWMnO
A surface response analysis was done using the conversion as
the response for the experimental plan. The p-values from this
analysis, where a p-valueo0.05 corresponds to a significant
factor, are displayed in Table 1. It is clear that the temperature
and gas flow rate had a significant effect on the conversion over
the response surface. A contour plot demonstrating the effect of
gas flow rate and temperature on conversion is shown in Fig. 5.
As expected, the conversion increased with an increase in
temperature. Somewhat unexpectedly however, was that the
conversion also increased with increased gas flow rate. An
increase in gas flow rate inherently causes the residence time to
AFR Product
Relative Intensity
3712
MnO Standard
Mn3O4 Standard
Mn2O3 Standard
20
30
40
50
2θ
60
70
80
Fig. 4. XRD product spectra from the AFR.
Table 1
Dissociation reaction response surface p-values.
Parameter
Constant
Temperature
Gas flow rate
X
0.00
0.05
0.02
MWMn2 O3
ð0:23O%Þ
1 MWO2
1 MWO2
1þ
þ 1þ
ðO%0:27Þ
4 MWMn3 O4
4 MWMnO
60
70
7
ð9Þ
Calculated conversions to MnO ranged between 50 and 74%, with
the remainder of the product consisting of Mn3O4.
Gas Flow (slpm)
0.35
0.3
0.25
Oxygen %
6
5
65
4
0.2
0.15
3
0.1
0.05
2
55
55
0
0
2
4
6
8
Time (min)
10
Fig. 3. O2 generation for Mn2O3 dissociation.
12
14
1650
1700
1750
1800
Temperature (K)
1850
Fig. 5. Contour plot of conversion to MnO.
1900
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decrease, which would typically lead to a decrease in conversion.
However, a high gas flow rate could be limiting the recombination/reverse reaction, where an increased bulk gas flow in the
system decreases the partial pressure of oxygen, which reduces
the potential for any released oxygen to find a surface site to
recombine with MnO to form Mn3O4 before leaving the system. In
addition, the speed at which the powder flows through the
quench tube is faster at higher flow rates, which in turn
diminishes the possibility of a recombination occurring in the
quench tube as well. In essence, there appeared to be a
recombination reaction occurring and the effect of the recombination is presented further in the next section.
3713
Recombination conversion to Mn3O4 as a function of O% was
derived in Appendix A and is
XMn3 O4 ¼ 3
O%0:23 MWMnO
MWMn3 O4
0:04
ð10Þ
Calculations indicated the conversion of MnO to Mn3O4 ranged
between 30 and 90%. Table 2 presents the p-value results of the
surface response analysis.
Unlike the dissociation, bulk gas flow rate did not have a
significant effect on conversion. The factors that did have an effect
on conversion were the concentration of oxygen and temperature.
Fig. 8 is a contour plot of oxygen concentration versus
T = 1270 O2 = 1.00%
The manganese oxide dissociation results discussed in Section 3.1
appeared to show that the recombination reaction was present to an
oftentimes significant degree. To investigate this, an experimental
plan using MnO as the reactant was employed. The oxygen trace from
a typical recombination reaction experimental run is shown in Fig. 6.
If the additional oxygen was not interacting with the feed powder, a
straight line would be expected.
From Fig. 6, as soon as the feed powder was introduced into
the reactor, the concentration of oxygen rapidly fell. After the
onset of the reaction, the oxygen concentration began to rise
again before two more distinct oxygen decreases. At the point of
these decreases light vibration had been applied to the particle
feeder to aid the flow of the powder. When the feed was shut-off,
a rapid increase in the amount of oxygen was observed, until it
leveled off at approximately the initial value. The observed
decrease in oxygen passing through the reactor was certainly
due to the recombination reaction occurring, as changes in the
oxygen concentration profile coincided with feed powder entering
the reactor. To confirm that the oxygen decrease seen in Fig. 6 was
actually from a recombination reaction, XRD was performed on
the samples that were collected from the gravity collection vessel.
Product XRD spectra for different temperature oxygen concentration combinations are shown in Fig. 7.
Comparing the product XRD spectra to the standardized
manganese oxide spectra indicated that the product was a
combination of Mn3O4 and MnO, without any Mn2O3. In addition,
a trend of increasing temperature and oxygen concentration
increased the amount of Mn3O4 in the sample. A surface response
analysis was performed on the calculated conversions to
investigate this trend statistically.
T = 1270 O2 = 0.25%
Relative Intensity
3.2. MnO to Mn3O4 recombination kinetics with oxygen
T = 700 O2 = 0.25%
MnO Standard
Mn3O4 Standard
Mn2O3 Standard
20
30
40
50
2Θ
60
70
80
Fig. 7. XRD product spectra from the recombination reaction.
Table 2
Recombination reaction response surface p-values.
Parameter
Constant
Temperature
Gas flow rate
Oxygen %
X
0.00
0.01
0.73
0.03
80
1.2
80
60
1.0
0.8
0.3
O%
0.25
Oxygen %
T = 700 O2 = 1.00%
0.6
0.2
0.4
0.15
40
0.1
70
0.2
0.05
0
2
4
6
8
Time (min)
10
Fig. 6. AFR O2 curve for recombination reaction.
12
14
50
30
0.0
0
800
900
50
1000 1100 1200 1300 1400 1500 1600 1700
Temperature (K)
Fig. 8. Contour plot of recombination conversion to Mn3O4.
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temperature with conversion as the response. The trend indicated
that as oxygen concentration and temperature were increased the
conversion of MnO to Mn3O4 increased. This result confirmed
what was observed from the XRD product spectra in Fig. 7.
It was particularly interesting that gas flow rate did not have
an effect on the conversion. What this result indicated was that
once the MnO was fed into the reactor, the recombination was
occurring almost instantaneously, and was limited based on
temperature and oxygen concentration. Additionally, the results
of the recombination experimental plan show that during the
quench of the full dissociation reaction, it is essential to maintain
a low oxygen concentration while cooling the dissociation
product as rapidly as possible to maximize the final Mn2O3 to
MnO conversion.
3.3. Mn2O3 to MnO forward reaction kinetics with oxygen
An experimental plan was executed to further understand the
effect of oxygen and temperature on the dissociation of Mn2O3.
Fig. 9 shows the oxygen generation curve normalized by the
amount of mass fed for the dissociation of manganese oxide with
and without oxygen added to the bulk gas (T¼ 1873 K, gas flow
rate¼7.0 slpm). It was fairly obvious by looking at the comparison
of the two curves, that the area under the curve without oxygen
0.2
With 1% Oxygen
Oxygen % / Mass Fed
0.18
0.16
0.14
0.12
0.1
0.08
Without Oxygen
0.06
was larger than with oxygen. What this shows is that a higher
conversion was attainable when oxygen was not present in the
bulk gas flow. The reaction products are further compared in
Fig. 10 using XRD spectra from each experimental condition.
The XRD spectra shown in Fig. 10 confirmed what was
observed from the oxygen generation curves. The main peak
from the MnO spectrum is more intense in the product spectrum
without oxygen in the bulk stream, suggesting a higher conversion.
Conversions to MnO, calculated using Eq. (9), were between 9
and 48%. Table 3 presents the analyzed results of the factorial
design.
The significant factors from the factorial design for Mn2O3
dissociation with O2 in the bulk gas flow were the same as those
from the recombination reaction. Both temperature and oxygen
percent had a significant effect on conversion, while bulk gas flow
rate did not.
Although the statistical results indicated that gas flow rate did
not have a significant effect on conversion, it showed an
interesting trend. When the bulk oxygen percent was at 0.25%,
the increase in gas flow rate caused an increase in conversion.
However, when the bulk oxygen percent was increased to 1.00%,
an increase in gas flow rate caused a decrease in conversion. This
result is evident when looking at XRD spectra of the reaction
products at 1673 K, as shown in Fig. 11.
When comparing the XRD spectra at an oxygen concentration
of 0.25%, it was evident that the main MnO peak (2y E411) at a
gas flow (GF) of 2.5 slpm was not visible. This peak became visible
when the bulk gas flow was increased to 7.0 slpm, which meant
that the conversion increased with increasing gas flow rate.
Conversely, when comparing the gas flow rates at a bulk O2
concentration of 1.00%, the Mn3O4 peak (2y E321) increased in
intensity when the bulk flow rate was increased from 2.5 to
7.0 slpm, equating to a conversion decrease with a gas flow
increase. Conversion calculations from experiments performed in
the AFR agreed with these observations and are plotted on the
design points in Fig. 12.
0.04
0.02
0
0
2
4
6
8
Time (min)
10
12
14
Fig. 9. O2 generation for Mn2O3 dissociation with and without O2 in the bulk gas
flow.
Table 3
Factorial design p-values for Mn2O3 dissociation with oxygen.
Parameter
Constant
Temperature
Gas flow rate
Oxygen %
X
0.00
0.00
0.56
0.05
GF = 7.0 O2 = 1.00%
Product With O2
GF = 7.0 O2 = 0.00%
Relative Intensity
Relative Intensity
GF = 7.0 O2 = 0.25%
Product Without O2
MnO Standard
GF = 2.5 O2 = 1.00%
GF = 2.5 O2 = 0.25%
GF = 2.5 O2 = 0.00%
MnO Standard
Mn3O4 Standard
Mn3O4 Standard
20
20
30
40
50
2θ
60
70
Fig. 10. XRD product spectra of Mn2O3 dissociation with and without O2.
80
30
40
50
2Θ
60
70
80
Fig. 11. XRD product spectra of Mn2O3 dissociation at different bulk O2
percentages at 1673 K (GF represents total gas flow rate).
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Fig. 12 confirms what was observed from XRD data. At low
oxygen concentrations, increasing gas flow rate increased
conversion. At higher oxygen concentrations, increasing gas flow
rate decreased conversion. An increased conversion with increased gas flow rate is what was observed during experiments
performed in an inert environment (Section 3.1). The results of
the experiments predict that as long as the oxygen concentration
is below 0.25%, the gas flow rate should have a significant effect
on conversion. Additionally, it shows that conversion trends from
two independent experimental plans were consistent.
The other trend that may be taken from Fig. 12 is that when
the reactor temperature was increased, the conversion increased;
when the oxygen concentration increased, the conversion
decreased. The trends are plotted in a contour plot in Fig. 13.
An interesting trend was observed when comparing the
contour plot in Fig. 13 with Fig. 8. In both contour plots, as the
oxygen concentration approaches 1% and temperature approaches
30.9
9.73
47.9
21.3
26.0
Gas Flow
(2.5-7 slpm)
3715
1700 K, both contours predicted a MnO fraction of approximately
20% and Mn3O4 fraction of approximately 80%.
When comparing the results of the experimental plans of the
recombination reaction with those from the forward reaction
with a set O2 percent in the bulk gas flow, it was apparent that a
limitation was being reached as the amount of oxygen was
increased to 1%. From the statistical model, if a significant amount
of oxygen is present in the system, the temperature must be
increased significantly in order to garner high enough conversions
for the Mn2O3 cycle to be feasible. Thus, it was concluded that for
the cycle to be feasible in an AFR, Mn2O3 must be dissociated in an
inert atmosphere.
3.4. Kinetic rate constant estimation from AFR data
There are many factors in the rapid dissociation of manganese
oxide that create difficulty in predicting the reaction kinetics of
Mn2O3 dissociation from the AFR data. The first difficulty is the
two step dissociation, with the second dissociation having a dual
mechanism. It was previously hypothesized that an Avrami–
Erofeev (Eq. (11)) nucleation and growth mechanism was
occurring from the beginning of the dissociation to the end of
the acceleration region of the second dissociation reaction
(Francis, 2008). Because this is a thermal reduction and the Biot
numbers are small, the Avrami–Erofeev mechanism predicts that
the reaction can take place anywhere in the particle if there is
sufficient energy.
39.9
11.9
RAvramiErofeev ¼ A1 eEa,1 =RT nð1XÞ½lnð1XÞðn1Þ=n
44.9
12.7
Oxygen
(0.25-1%)
Temperature
(1673-1873 K)
Fig. 12. Conversions from Mn2O3 dissociation with oxygen.
Because the same reaction mechanism was occurring throughout
the entirety of the Mn2O3 to Mn3O4 reaction, and the first portion
of the Mn3O4 to MnO reaction, both of these mechanisms were
grouped into a single mechanism when estimating a rate
constant. The higher activation energy from the second dissociation Avrami–Erofeev mechanism was used in the calculation, to
ensure a conservative estimate of the reaction rate constant.
Therefore the pre-exponential rate constant was calculated (for
conversions that were equal or less than 60%) using Eq. (12).
1.0
A¼
20
ð11Þ
ðlnð1XÞÞ1=n eEa=RT
t
ð12Þ
The second issue that arose was that when the conversion
exceeded 60%, the total degrees of freedom did not equal zero, as
the mechanism shifted to an order of reaction mechanism:
0.9
0.8
RMn2 O3 ,2 ¼ A2 eEa,2 =RT ð1XÞn
Oxygen %
0.7
0.6
40
0.5
To overcome this issue, the rate constant was calculated for all
runs that had conversions equal to or less than 60%, which was
approximately half of the experimental runs. For the runs that
exceeded 60%, the time for the reaction to attain a conversion of
60% was calculated by rearranging Eq. (12). This time was then
subtracted from the total residence time, to calculate the time in
which the order of reaction mechanism controlled the reaction:
t2 ¼ t
0.4
1700
1750
1800
Temperature (K)
ðlnð0:4ÞÞ1=n eEa =RT
A
ð14Þ
To then calculate the rate constant, the order of reaction
mechanism was integrated from 0.6 to X, and subsequently
divided by the residence time calculated from Eq. (14):
30
0.3
ð13Þ
1850
Fig. 13. Contour plot of dissociation conversion to MnO with oxygen.
ð1XÞ1n ð0:4Þ1n Ea =RT
e
n1
A2 ¼
t2
ð15Þ
ARTICLE IN PRESS
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T.M. Francis et al. / Chemical Engineering Science 65 (2010) 3709–3717
aspects of the industrial scale-up of the high temperature
reduction step of metal oxide water-splitting cycles.
Table 4
Predicted rate constants for Mn2O3 dissociation.
TGA
AFR
A1 (s 1)
A2 (s 1)
5.8 104 7 1.5 103
1.8 107 7 1.3 107
5.8 102 7 8.4 102
5.6 103 7 4.1 103
The kinetic rate parameters that were calculated in this section
were based on the experimental plan that was performed in an
inert environment as the oxygen was dilute enough to not
significantly affect the reaction. Additionally, the reaction
mechanisms were determined in an inert atmosphere. With that,
approximately half of the runs yielded conversions of less than
0.6. The calculated rate constants are presented in Table 4. Values
calculated from a TGA are included for comparison.
The confidence intervals reported in Table 4 for the reaction
kinetics parameters are rather large, but are typical for AFR
analyses (Himmelblau, 1970; Galwey and Brown, 1999; Carney
et al., 2006a, b; Perkins et al., 2008). Qualitatively, the reaction
rates from the AFR were three orders of magnitude faster than
from a TGA. This increase in reaction rate was believed to result
from rapid oxygen diffusion away from dispersed particles in the
AFR. The results indicate that an AFR is an excellent choice of
equipment for the dissociation of manganese oxide. The increased
reaction rate and high overall conversion are especially important
for the feasibility of the sodium manganese thermochemical
water-splitting cycle.
4. Conclusions
Manganese oxide dissociation was investigated in an AFR,
because of its importance as part of a thermochemical watersplitting cycle for the production of hydrogen using renewable
means. The investigation showed that high Mn2O3 conversions
are achievable with the use of an AFR. This is significant for
implementation and scale up of a Mn2O3-based thermochemical
cycle. Additionally, the results showed that by using an AFR, the
reaction is rapid, with kinetic rate constants that are orders of
magnitude greater than what was observed using a TGA.
Overall conversions to MnO ranged between 50 and 74%,
depending on the temperature and gas flow rate. The highest
conversion was achieved with a combination of a high temperature and high bulk gas flow rate. It was clear that an increased
temperature favored the forward reaction. Conversion remained
high when high gas flow rates were used. This gave validity to the
hypothesis that the propensity for the recombination reaction to
occur could be effectively suppressed using high gas flow rates.
The effect of oxygen on the dissociation was also investigated.
When the oxygen concentration exceeded a critical point (0.25%),
the recombination reaction apparently dominated the reaction
and decreased the net conversion to MnO. The only way to obtain
high conversions at oxygen contents exceeding 0.25% was to
increase the temperature. Below the critical oxygen concentration, however, gas flow rate and temperature controlled the
dissociation. What was especially interesting was that increasing
gas flow rate, essentially decreasing residence time, increased
conversion. This was again hypothesized to occur because of the
faster quench step the reaction was undergoing, limiting the
potential of the recombination reaction from occurring. Together,
the results of these studies lead to the conclusion that high
conversion was possible provided that the dissociation can be
performed in an inert gas environment with a rapid quench. This
work provides insight into the dissociation kinetics of Mn2O3, and
Notation
MWMn2 O3
MWMn3 O4
MWMnO
MWO2
MWMn
Msample
NFe
NFe,ex
NKMnO4
NMn
NMn3 þ
NMn2 þ
NMn2 O3
NMn3 O4
NMnO
O%
XMn3 O4
XMnO
YMn3 O4
YO2 ,Mn3 O4
YMnO
YO2 ,MnO
molar mass of Mn2O3, g/mole
molar mass of Mn3O4, g/mole
molar mass of MnO, g/mole
molar mass of O2, g/mole
molar mass of O2, g/mole
sample mass, g
number of Fe2 + moles, moles
number of excess Fe2 + moles, moles
number of KMnO4 moles, moles
moles of Mn, moles
moles of Mn3 + , moles
moles of Mn2 + , moles
moles of Mn2O3, moles
moles of Mn3O4, moles
moles of MnO, moles
oxygen percent from LECO TC-600, unitless
conversion to Mn3O4, unitless
conversion to MnO, unitless
mass fraction of Mn3O4, unitless
mass fraction of O2 from Mn3O4 transition, unitless
mass fraction of MnO, unitless
mass fraction of O2 from MnO transition, unitless
Acknowledgments
The authors wish to thank the United States Department of
Energy (DOE) Grant number DE-F636-036013062 and the UNLV
foundation for funding this project. Additionally, the authors would
like to thank N. Brandon Hughes for the titration work. One of the
authors wishes to acknowledge the GEM Consortium Fellowship
program, National Science Foundation AGEP Fellowship program,
and the United States Department of Education GAANN program
for providing partial research scholarships.
Appendix A
A.1. Determining forward conversion with oxygen analysis
In calculating the conversion of Mn2O3 to MnO, it was assumed
that Mn2O3 did not exist in the sample. XRD patterns supported
this assumption as there were only peaks attributed to the
presence of Mn3O4 and MnO. A mass balance shows that when
starting with 100% Mn2O3 there can only be Mn3O4, MnO, and
oxygen released from the two reactions. Dividing by the initial
mass of Mn2O3 the mass balance is shown in Eq. (A.1):
1 ¼ YMn3 O4 þ YO2 ,Mn3 O4 þ YMnO þYO2 ,MnO
ðA:1Þ
Oxygen percent, found using a LECO corp. TC-600 Oxygen
Determinator, was put in terms of mass fraction of Mn3O4 and
MnO (Eq. (A.2)).
O% ¼
0:27YMn3 O4 þ 0:23YMnO
YMn3 O4 þ YMnO
ðA:2Þ
Oxygen fractions of Mn3O4 and MnO were 0.27 and 0.23,
respectively. To find the mass fraction of O2 released during the
transition to Mn3O4, an overall mole balance of the dissociation
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T.M. Francis et al. / Chemical Engineering Science 65 (2010) 3709–3717
Rearranging Eq. (A.12) leads to the mass fraction of Mn3O4 in the
product:
was performed:
6Mn2O3-4Mn3O4 + O2-12MnO+ 2O2
(A.3)
YMn3 O4 ¼
From Eq. (A.3), for every four moles of Mn3O4 there will be one
mole of O2 contributing to this fraction; for every twelve moles of
MnO there will be on further mole of O2. Putting the mole balance
in terms of the contributions from Mn3O4 and MnO, the mass
fractions yields
YO2 ,Mn3 O4
YMn3 O4 MWO2
YMnO MWO2
¼
þ
4
MWMn3 O4
12 MWMnO
ðA:4Þ
A similar balance for the mass fraction of O2 released from the
transition to MnO yields
YO2 ,MnO ¼
YMnO MWO2
6 MWMnO
ðA:5Þ
Placing Eqs. (A.4) and (A.5) into Eq. (A.1) and simplifying, yields
Eq. (A.6):
1 MWO2
1 MWO2
þ YMnO 1 þ
ðA:6Þ
1 ¼ YMn3 O4 1 þ
4 MWMn3 O4
4 MWMnO
Eqs. (A.2) and (A.6) provided two equations with two unknowns
resulting in zero degrees of freedom. Solving Eq. (A.2) for YMn3 O4
and substituting into Eq. (A.6) yields
YMnO ð0:23O%Þ
1 MWO2
1 MWO2
þ YMnO 1 þ
1þ
1¼
ðO%0:27Þ
4 MWMn3 O4
4 MWMnO
ðA:7Þ
Solving for YMnO gave Eq. (A.8):
YMnO ¼
1
ð0:23O%Þ
1 MWO2
1 MWO2
þ 1þ
1þ
ðO%0:27Þ
4 MWMn3 O4
4 MWMnO
ðA:8Þ
Assuming 1 mole of Mn2O3 was reacted, conversion to MnO may
be defined as
XMnO ¼
YMnO MWMn2 O3
2
MWMnO
ðA:9Þ
Substituting Eq. (A.8) into Eq. (A.9) gave conversion to MnO as a
function of O%:
XMnO ¼
MWMn2 O3
ð0:23O%Þ
1 MWO2
1 MWO2
1þ
2MWMnO
þ 1þ
ðO%0:27Þ
4 MWMn3 O4
4 MWMnO
ðA:10Þ
A.2. Determining recombination conversion with oxygen analysis
To calculate the conversion for the recombination of MnO to
Mn3O4, only MnO and Mn3O4 need to be included, as verified by
XRD. Eq. (A.2) was used again to relate the mass fractions of MnO
and Mn3O4 to the amount of oxygen in the sample. The mass
balance this time consists of MnO and Mn3O4:
1 ¼ YMn3 O4 þYMnO
ðA:11Þ
Eq. (A.11) was rearranged and substituted into Eq. (A.2):
O% ¼ 0:27YMn3 O4 0:23YMn3 O4 0:23
3717
ðA:12Þ
O%0:23
0:04
ðA:13Þ
Assuming 1 mole of MnO was reacted conversion to Mn3O4 may
be defined as
XMn3 O4 ¼ 3YMn3 O4
MWMnO
MWMn3 O4
ðA:14Þ
Substituting Eq. (A.13) into Eq. (A.14) gave conversion for the
recombination from MnO to Mn3O4 as a function of O%.
XMn3 O4 ¼ 3
O%0:23 MWMnO
MWMn3 O4
0:04
ðA:15Þ
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