EPA-AA-TSS-84-7
Technical Report
Methanol Dissociation: The Effects
of Partial Dissociation
By
William B. Clemmens
December 1984
NOTICE
Technical Reports do not necessarily represent final EPA
decisions or positions. They are intended to present
technical analysis of issues using data which are
currently available. The purpose in the release of such
reports is to facilitate the exchange of technical
information and to inform the public of technical
developments which may form the basis for a final EPA
decision, position or regulatory action.
Technical Support Staff
Emission Control Technology Division
Office of Mobile Source Air Pollution Control
Office of Air, Noise and Radiation
U. S. Environmental Protection Agency
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EPA-AA-TSS-84-7
Technical Report
Methanol Dissociation: The Effects
of Partial Dissociation
By
William B. Clenunens
December 1984
NOTICE
Technical Reports do not necessarily represent final EPA
decisions or positions. They are intended to present
technical analysis of issues using data which are
currently available. The purpose in the release of such
reports is to facilitate the exchange of technical
information and to inform the public of technical
developments which may form the basis for a final EPA
decision, position or regulatory action.
Technical Support Staff
Emission Control Technology Division
Office of Mobile Source Air Pollution Control
Office of Air, Noise and Radiation
U. S. Environmental Protection Agency
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-2-
Table of Contents
Page
I. Introduction 3
II. Summary 4
A. Procedure, Results, and Observations 4
B. Conclusions 14
* * * * Supporting Analyses and Technical Information * * * *
III. Test Set-Up 16
IV. Engine Operation 17
V. Results
A. Efficiency, Load 1 17
B. Efficiency, Load 2 19
C. Emissions 19
VI. Discussion
A. Brake vs. System Efficiency 24
B. Lean Limit 27
C. Minimum Required Dissociation 33
VII. Postscript 36
References 38
Appendix I Stoichiometric Fuel-Air Ratios 40
Appendix II Equivalence Ratio at the Lower
Flammability Limit 46
Appendix III Energy Values and Derivations 48
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Methanol Dissociation: The Effects
of Partial Dissociation
I. Introduction
The Office of Mobile Sources within the Environmental
Protection Agency has studied and evaluated alternative
transportation fuels since its formation in 1970. EPA's
responsibilities under the Clean Air Act also have
necessitated a significant regulatory role dealing with
transportation fuels. In particular, Section 211 of the
Clean Air Act requires EPA to play a key role in the
introduction of new fuels and fuel additives. Perhaps most
visible was EPA's role in the introduction of unleaded
gasoline to permit the use of catalytic converters on 1975
and later model year automobiles. More recently EPA has
responded to a growing interest in the use of oxygenates (in
particular methanol) for use in motor vehicles and for
blending with gasoline.
As part of this response to the interest of methanol, a
program was undertaken to grade several fuel utilization
concepts. One of those concepts that has been evaluated is
what we have termed "partial dissociation." The process of
dissociation is the break-up of one chemical combination into
simpler constitutents. In the case of methanol fuel, the
combination of CH3OH is broken down with heat and a
catalyst into a gaseous mixture of approximately 66 mole
percent H2 and 34 mole percent CO. The theoretical
advantages of utilizing dissociated methanol over liquid
methanol for engine fuel are two-fold. One, the energy of
the dissociated mixture is approximately 10,000 BTU/lb versus
8,644 BTU/lb for liquid methanol. Potentially with the
higher BTU content of the dissociated fuel, each gallon of
liquid methanol can be utilized much more efficiently,because
the heat needed to dissociate the liquid methanol can be
obtained from the waste heat of the engine exhaust gas. And
two, the nature of hydrogen combustion should allow engine
calibrations that provide leaner operation which can imply in
some cases better efficiency and lower emissions. Combined
the two factors (enhanced energy content and hydrogen
combustion) have the potential to improve the thermal
efficiency by about one-fourth over and above the base
efficiency, if the fuel is completely dissociated (see
Section II.A). A maximum of about 16 percent improvement
over the base can be derived from the energy enhancement
alone. The balance of the improvement relates to how well an
individual engine utilizes the enleanment allowed by the
hydrogen addition.
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Typically, however, the volume requirements for the catalyst
and the heat exchanger necessary to dissociate all of the
methanol used by the engine are quite large leading to
vehicle space limitation problems. Also, the carburetor
control system for large quantities of H2/CO gas can be
complex. Because the theoretical possibilities looked
promising, the question was posed, "Is it possible to achieve
a sufficient portion of the theoretical potential of full
dissociation with some lower amount of dissociation and
correspondingly less complex dissociation hardware?" In
other words, can more space-efficient part load performance
be obtained by splitting the energy flow to the engine, the
major fraction being liquid methanol and the smaller fraction
being passed through a compact dissociation reactor (with
possibly better dissociation efficiencies), and could this
combination result in a better overall package than
attempting to dissociate the entire fuel flow in a large
reactor which typically exhibits relatively poor dissociation
efficiencies at part load?
II. Summary
The results indicate that possibly as much as 40 percent of
the potentially available improvements can be obtained
utilizing only a 30 percent fraction of the simulated
dissociation gas. However, the absolute magnitude of the
improvement in efficiency at these fractional dissociation
levels was somewhat limited. The method of energy accounting
is critical to the level reported up to 5 percent
improvement by one technique and up to 10 percent improvement
by the other. Also there appears to be a minimum level of
dissociation necessary to obtain positive improvements 'in
efficiency. This level coincided with improvements in
emissions.
Highlights and observations of these results as well as
conclusions drawn from them are presented in this section.
Supporting analyses and technical information are contained
in the following sections and appendices.
A. Procedure, Results, and Observations
To test the hypothesis, a series of tests were run on a
Nissan 2 liter NAPS-Z engine converted to methanol
operation. A mixture of hydrogen and carbon monoxide
(H2/CO) gas was injected into the intake manifold from
cylinders of compressed gas to simulate low levels of
dissociation (Figure 1). The amount of H2/CO gas injected
was approximately 4, 9, 18, and 30 percent by mass of the
total energy flow to the engine. Note percent dissociation
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AIR FILTER
. FUMIGATION
INJECTORS
(LOCATION 2)
I 0 0
P N
FUMIGATION
TEST SET-UP
FIG 1
Table II-l
Simulated Dissociation at
1500 RPM and 29.5 ft-lb
Simulated Dissociation
Weight Fractions
Baseline 4%
9t
18%
30%
% Dissociation
BTU basis
Mass basis
5.5
4.5
11.6
9. I
26.4
18.5
S3. 4
31.4
Equivalence Ratip
(for best results)
.69
.647
.701
.618
.583
Best Efficiency
Thermal
Efficiency (% BTE) 26.6 26.63 27.08 28.02 27.82
System
Efficiency (% SYE) 26.6 26.58 27.44 28.88 29.29
Difference 0 0 .4 .9 1.5
\ Improvement *
BTE
SYE
1.8%
3.2%
5.3%
8.6%
4 6% ,'n
10.1%
'Compared to neat methanol baseline
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is used in this paper to describe the mass fraction of the
total energy consumed by the engine which was H2/CO gas.
(The total energy was assumed to be an equivalent amount of
liquid methanol which implies a dissociation efficiency of
100 percent in a small reactor.)
Because of the complexity of the energy accounting, it was
important to identify the two accounting techniques used to
determine thermal efficiency. The distinction between the
techniques is where the boundaries of the control volume are
placed. Brake Thermal Efficiency (BTE) was calculated on an
energy-out over energy-in basis (i.e. fuel entering the
control volume boundary and power exiting). In our
simulation the fuel crossing the boundary was a mix of
methanol, H2 , and CO. However, if a dissociation reactor
had been used, the only fuel crossing the control volume
boundary would be methanol. The concept of System Thermal
Efficiency (SYE) was developed to transform our measured BTE
data into results that might be expected as an upper limit of
performance (SYE) if a reactor had been used (an analytical
description of this treatment is found in section VI-A).
With these conventions firmly in mind (BTE vs. SYE), the test
results in Table II-l indicate that modest improvements in
brake thermal efficiency (BTE) were possible (on the order of
5 percent improvement). Figure 3* shows the trends of these
results while Figure 2 indicates the baseline curve.
However, system efficiency (SYE) improvements of up to 10
percent are calculated for the highest partial dissociation
rate simulated, if the H2/CO gas were assumed to come from
liquid methanol, and if the conversion from liquid methanol
through vapor phase to dissociation gas were complete (see
Table II-l and Figure 4).
To put these data in perspective, we compared our partial
dissociation data to data on full dissociation (100% of total
energy flow) reported by Hirota (Ref. 1).** Because Hirota
used a true dissociation reactor, his results are in units of
BTE. In order to compare our results to his, our SYE results
* Note: The Greek work "phi" is used to designate equivalence
ratio throughout the report.
**Numbers in parentheses designate References at the end of
paper. In Hirota's study(l), a dissociation reactor was used
to provide the dissociation gas. Analysis of the gas
indicated a mixture of components, the dominate species being
H2 and CO. For the purposes of comparison, it was assumed
this mixture represented complete dissociation of the total
engine fuel flow.
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BASELINE BRAKE THERMAL iff I CIENCT IBTEI
29
28
27
26
25
"
23
22
21
20
/
/ S
1500 RPM
29.5 ft-lb .
X BASELINE (NEAT
NETHANOL. 13:1 CHI
N-7
S STOCK ICASOLINE. 9: 1
CRI
N.I
O.itS O.SS 0.65 0.75 0.85 0.95 1.05
0.50 0.60 0.70 0.80 0.90 1.00
PHI
FIG 2
BRAKE THERMAL EFFICIENCT U/PARTIAL DISSOCIATION
10
29
28
27
26
£ 25
CD
2>4
23
22
21
20
»
L
^\*
£*^1 ^
/ ^~~"*~~~T~
/
i
f
/ 5
Q
a
1500 RPM
29.5 ft-lb
, , , ,
X BASELINE INERT
METHANOLI
N-7
I8/. HZ/CO BT MASS
N-S
0 302 H2/CO 8T MASS
N-S
S STOCK (GASOLINE. 9: 1
cm
N.I
0.i*S 0.55 0.6S 0.75 0.65 0.95 I.OS
0.50 0.60 0.70 0.80 0.90 1.00
PHI
FIG 3
STSTEN TMEBMRL EFMCIENCT H/ PflRTIBL OISJOCIflTION
30 ,
29
28
27
26
23
22
21
20
O.US O.SS 0.65 0.75 0.85 0.95 1.05
0.50 0.60 0.70 0.80 0.90 1.00
PHI
1500 RPM
29.5 ft-lb
X BASELINE (NEST
METMANOL)
N-6
1BX H2/CO 8T MASS
N-5
0 307. H2/CO 8T MAS]
N-S
FIG 4
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STSTEM THERMflL EFF I C IENCY (STE) VS PERCENT DISSOCIRTION
UJ
35
34
33
32
31
30
29
28
27
26
25
24
23
22
21
20
EPA Data
1500 RPM
29.5 ft-lb
Hirota's Datad
1600 RPM
28.9 ft-lb
i i
X PHI-.583 -.588
N-3
Y PHI-.6H7 -.667
20 40 60 80 100
10 30 50 70 90
PERCENT H2/CO BY MflSS
FIG 5
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(which simulated the use of a reactor with optimum conversion
efficiency) were compared to Hirota's BTE results. Figure 5
is a plot of this data (efficiency) versus the fraction of
dissociation gas used. There is a great deal of
interpolation between 30 percent dissociation (our highest
limit) and 100 percent dissociation (Hirota's data) in Figure
5.* However, these limited results do suggest that there
seems to be a greater fraction of improvement at the lower
levels of dissociation than at the higher levels.
This observation is supported by several other factors
involved in the dissociation process. First to digress a
little, the source of the improvement in efficiency from the
dissociation process appears to eminate from two separate
processes. The initial process occurs by the introduction of
hydrogen into the fuel. The amount of improvement seems to
be tied to the ability of a particular engine to run leaner
with less cycle to cycle variation when hydrogen is added
than without the hydrogen addition. (The improvement in
cyclic variation was readily apparent by subjectively
observing the engine roughness during testing). To some
extent this ability is governed by the lean limit. A review
of the literature (see section VI-B) suggests that the
improvement in the lean limit can be correlated with the
level of hydrogen added to the fuel (HWF - hydrogen weight
fraction), and suggests that the percentage improvement in
the lean limit is greater at the lower to moderate levels of
hydrogen addition. The second process involves the energy
enhancement of the fuel resulting from the dissociation
process. The maximum improvement from this source can be
calculated from the lower heating value of the composite
gases used which in our case was for a hydrogen and carbon
monoxide mixture. Using the lower heating values (from
Table III-l), we find that the value for our mixture of
hydrogen and carbon monoxide is 16 percent higher than the
value for liquid methanol. The improvement for the second
process is therefore limited to 16 percent at 100 percent
dissociation (see section VI-A). This maximum limit is, of
course, proportioned based on the fraction of partial
.dissociation. The observed data suggest that this
relationship also is not linear, and that a greater fraction
of enhancement occurs at the lower rates of hydrogen addition
to the fuel. (This may be due to the relationship between
the hydrogen weight fraction (HWF) and the hydrogen energy
fraction (HEF) - see Figure 12, section VI-B.)
*In most cases, the data plotted in this report represent the
average of multiple results. However, in some cases single
data points were used. In general, the validity of those
points is subject to how well they fit the trends of the
other data.
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Both sources of efficiency improvement from dissociating
methanol, then, suggest that a greater fraction of
improvement would be expected to occur at lower, or partial
dissociation rates, and supports the similar observation
determined from the test data. From this we can conclude
that while the hypothesis appears to be basically true (more
change at low dissociation rates) , the potential improve-
ments in system efficiency are not exceedingly large at the
low rates - typically less than 10 percent improvements. If,
however, we assume Hirota's (1) base engine on neat methanol*
would have similar efficiency to our base engine (which is
implied in Figure 5), full dissociation only nets around a 26
percent improvement over the base. From this aspect, our 10
percent improvement in SYE represents about 40 percent of the
available improvement. An 8 percent SYE improvement would
represent about 30 percent of the available increase. Since
these improvements are approximately the results obtained for
the 30 percent fraction and the 18 percent fraction
respectively, we see that the return on efficiency at these
lower rates may possibly exceed the fraction of dissociation
necessary for that return. This perspective may justify the
potentially less complicated hardware necessary for partial
dissociation under certain conditions.
Another interesting aspect of the partial dissociation
exercise is that the introduction of the hydrogen gas in the
Hz/CO mixture might be considered an approximate method to
simulate the effects of extending the lean misfire limit by
some other means. If this assumption is accepted, then we
might consider the brake thermal efficiency (BTE) (as opposed
to system efficiency) with the H2/CO gas as somewhat
indicative of the BTE that could be achieved with liquid
methanol if the lean limit were somehow extended by other
means. Figure 3 indicates the baseline BTE performance with
liquid neat methanol and the higher levels of H2/CO
addition as the equivalence ratio is changed. In the
baseline case, leaning the engine down to about an
equivalence ratio of about 0.69 (from about 0.85), improves
the efficiency slightly; below 0.69 the BTE falls off
dramatically (assumed to be the result of increasing misfires
as evidenced by the increasing organic matter emissions in
Figure (6). As evidenced from the 4 percent and 9 percent
*Hirota (1) did not provide data in his paper on any baseline
operation on neat methanol. However, since he used a similar
engine, a Nissan L20 versus a Nissan Z20, we felt such an
assumption was justifiable to estimate the return from
partial dissociation even though the estimate is sensitive to
the baseline level.
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data (reported elsewhere in this paper), doping the engine
with small amounts of H2/CO gas causes the improvement in
BTE performance to continue somewhat below the 0.69 level,
but then the BTE falls off as in the baseline case.
Increasing the amount of H2/CO gas to 18 percent and 30
percent appears to extend this fall-off point, and the BTE
performance improvement was continued down to an equivalence
ratio of 0.58. At this ratio, lean misfire was encountered
even though the equivalence ratio was still considerably
above the 0.45 lean flamability limit of methanol. However,
prior to reaching the abrupt lean misfire limit, no fall-off
in BTE performance was encountered, nor were any sudden
increases in organic matter (OM) emissions evident (which
appears to be indicative of BTE fall-off). It is also
interesting to note that between an equivalence ratio of 0.70
to 0.75, the BTE performance of liquid methanol and partially
dissociated methanol is for all practical purposes the same.
If the preceeding scenario is an appropriate representation
of the hydrogen addition, then the partial dissociation gases
could be construed as an "ignition enhancement device" for
lean methanol operation. Some support for this perception is
reported by Inagaki, et al. (13) suggesting that the minimum
ignition energy for reformed or dissociated gas is on the
order of 10 times less than that needed for gasoline.
Inagaki did not report a comparison with neat methanol, but
such a comparison would be expected to be of the same order
of magnitude in that the required ignition energy between
gasoline and methanol is generally accepted to be about the
same (typically 0.25 MJ for gasoline and 0.23 MJ for
methanol). Therefore, it seems logical that the fractional
dissociation would enhance the initial rate and the
completeness of combustion through 'an improved ignition
reserve potential relative to neat methanol. Both factors
(combustion rate and completeness) are generally believed to
be advantageous to lean engine operation with good efficiency.
Complicating the effectiveness of the potential ignition
enhancement of partial dissociation is that a minimum amount
of dissociation seems to be required to obtain efficiency
improvements with lean mixtures (details are in section
VI-C). Otherwise efficiency losses (from optimum with 100%
liquid methanol) occur. This minimum required dissociation
(MRD) level appears to be around 12 or 13 percent
dissociation as a fraction of the total fuel flow at the
primary test point, (1500 RPM, 29.5 ft-lb) . At a lower load
point the MRD level seemed to increase to between the 20
percent and 34 percent dissociation levels (no test points
were run between these values to narrow the increment).
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OM EMISSIONS W/PBBTIflL DISSOCIATION
1500 RPM
29.5 ft-lb
.40 .US .SO . 55 .60 .65 .70 .75 .80 .85 .90
X BflSELINE
N.JM
<4X H2/CO BT MflSS
N.a
9X H2/CO BT MflSS
N- I I
0 IBX M2/CO BT MUSS
N- IB
CO EMISSIONS H/P«RT1PL OI550CIBTION
BASELINE
1500 RPM
29.5 ft-lb
X BASELINE
N.3M
T/ M2/CO BT MflSS
N.8
9X M2/CO BT MBSS
N-l 1
0 ISX H2/CO BT MflSS
N-18
FIG 6
.HO .MS .50 .55 .60 .65 .70 .75 .80 .85 .90
PMl
FIG 7
M
I
NOX EMISSIONS M/PflBTIBL DISSOCIATION
1500 RPM
29.5 ft-lb
.10 .>45 .50 .55 .60 .65 . 70 . 75 .80
X BASELINE
N-3M
MX M2/CO BT MflSS
N.8
9Z M2/CO BT MflSS
N-l I
0 IBX H2/CO BT MflSS
N-ie
as . 90
FIG 8
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Engine emission results from a limited series of tests are
presented in Figures 6, 7, and 8 plotted against the actual
equivalence ratio. OM* and CO tend to follow the classic
lean misfire trends. They seem to be affected by the level
or fraction of dissociation gas used. The trend in HC and CO
at the 18 percent level are what one might expect from
improved burn rate, and this improved emission performance
corresponds with the improved efficiency that appeared to
occur above the MRD level. Therefore, the emission data
would tend to support the hypothesis that a minimum level of
dissociation is required for an efficiency improvement. NOx
emissions seem to be primarily dependent only on overall
equivalence ratio.
Since both the minimum dissociation (MRD) level and emission
results tend to support an ignition enhancement role for
small fractions of dissociation, it seems reasonable to pose
the question, "are there other means of ignition enhancement
that are as effective, less costly, less complex, and
possibly more predictable?" Improvements to the standard
ignition system are a first thought, and we have seen some
slight improvement with this approach. Of course, plasma
ignitors have generally been tried by various researchers on
gasoline fueled SI engines with mixed results. One
possibility that apparently hasn't been tried is to move the
source of the dissociation gas closer to the spark plug. In
other words, create a locally rich mixture of dissociation
gas near the spark plug at the time of ignition, rather than
fill the whole chamber with dissociation gas (i.e., the
stratified approach could reduce the amount of dissociation
gas needed). Three possibilities to achieve this stratified
ignition mixture are (1) deliver the dissociation gas
directly to the plug or the vicinity of the plug (possibly in
a prechamber), (2) utilize some dissociation catalyst
material within the spark plug (a plasma plug might be more
appropriate), or (3) dissociate methanol in the chamber near
the spark plug.
The ability of any of these (or any other) approaches to
provide sufficient dissociation gas necessary to exceed the
MRD level is unknown. However, future work on either
extending the lean limit or evaluating fractional
dissociation should consider the following conclusions.
*0rganic Matter (OM) values were determined by a standard LDV
FID. The OM values calculated include the weight of oxygen
in the hydrocarbon density of the exhaust. The density value
used was 37.69 g/cu ft (vs. 16.33 g/cu ft for gasoline HC) .
No correction for the FID response to methanol was applied.
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B. Conclusions
1. Future work with dissociation should use weight
fractions of the components for comparison work.
It was found to be particularly useful to separate
the hydrogen weight fraction (HWF) for lean limit
correlations
2.
Efficiency values can change based on the energy
accounting technique used. In general, it is
recommended that the actual results be the primary
units and simulated results be secondary, however,
ecific techniques used should be described.
the speci
3. Moderate improvements in brake efficiency are
possible with small fractions of methanol
dissociation. Improvements up to 10 percent
(simulated - 5 percent actual) were observed with
a 30 percent fraction.
4. There appears to be a threshold or minimum
required dissociation (MRD) level that is
necessary before improvements can be obtained with
lean mixtures. This hypothesized MRD level may
change with speed and load.
5. OM and CO emission levels seem to be influenced by
the MRD level. NOx seems only dependent on
equivalence ratio.
6. A dissociation reactor has the potential to
effec--vely enhance the energy content of the fuel
by about 16 percent above that obtained from lean
operation, but only if all of the fuel reaching
the engine is dissociated, and consists of
predominately hydrogen and carbon-monoxide in
their equilibrium proportions. The potential
improvement would be correspondingly less if the
dissociation process were not complete. For our
fraction of dissociation with equilibrium
concentrations, the maximum enhancement would be
expected to be on the order of 5 percent.
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Supporting Analyses
and
Technical Information
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III. Test Set-Up
The engine used for this test series was a standard 2 liter
Nissan NAPS-Z engine (with electronic port fuel injection
-EFI) modified for methanol use. The compression ratio was
12.8:1. Minimum timing for best torque (MET) was used
throughout the test series,* and exhaust gas recirculation
was eliminated.
The hydrogen and carbon monoxide (H2/CO) gas mixture was
obtained from compressed gas cylinders and the cylinders were
gravimetrically analyzed (±1 percent) to assure an accurate
indication of H2/CO split. The heating value of the
mixture was then calculated on this split based on the
heating values of the pure components. (See Table III-l, for
more details see Appendix III.)
Table III-l
Net Heating Values**
of Fuels Used
Compound Net Value (LHV) (BTU/lb)
H2 51,743
CO 4,328
Mix 10,039
Methanol 8,644
**see Appendix III
The H2/CO mixture entered the engine through the passage
normally used for the hot idle compensator circuit (see
Figure 1) . This location is slightly upstream of the EFI
throttle valve and between the throttle valve and the EFI air
meter. An alternate location was briefly attempted, but time
pressures prevented a thorough investigation. The alternate
location was through the stock EGR entrance to the inlet
manifold.
Because of the quantities of H2/CO gas mixture required to
achieve even moderate dissociation rates, only lower power
*MBT was used throughout the series when attainable; in many
cases timing was not advanced past incipient detonation, even
though more advance would have been required to achieve true
MBT.
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points were considered for investigation. Since technical
personnel from one automobile manufacturer had previously
suggested that a test point of 38 psi BMEP at 1500 RPM was
representative of overall in-vehicle engine performance, we
selected that point as one of our two test conditions. For
this engine, 38 psi BMEP and 1500 RPM works out to be
approximately 8 BHP or approximately 29.5 ft. Ib. of torque
(Load 1). The second test condition was arbitrarily selected
at 10 ft. Ib. (Load 2) and 1500 RPM to give us potentially
greater flexibility to achieve higher dissociation rates
(because of limited H2/CO flow rate capability).
Emission measurements were dilute bag measurements. A
standard light-duty vehicle (LDV) critical flow venturi (CFV)
CVS was used. However, due to the increase in water vapor
from the methanol engine, a refrigeration trap was installed
in the sample line between the probe and the sample bags.
The sample bags were analyzed with a standard LDV analytical
system which included a non-heated FID HC analyzer.
IV. Engine Operation
Engine operation with the H2/CO gas mixture did not differ
in any quantitative way from operation with neat methanol.
From a qualitative aspect the only observation apparent was
that the engine seemed to run smoother with the H2/CO gas
than neat fuel between an equivalence ratio of about 0.7 to
about 0.58. Below 0.58 overall equivalence ratio, the engine
would not consistently operate smoothly, regardless of
dissociation rate utilized (The highest rate attempted was 30
percent at 29.5 ft-lb).
V. Results
A. Efficiency, Load 1
The testing with H2/CO was performed in two series of
tests. The first series was tested at the 29.5 ft-lb
condition with fixed H2/CO flow rates of 15, 30, and 60
SCFH. These flow rates correspond to dissociation rates of
approximately 4, 9, and 18 percent by mass of the neat
methanol fuel flow in the baseline condition for the range of
equivalence ratios tested (i.e., there was sufficient
latitude to adjust the liquid methanol flow rates to achieve
various equivalence ratios at a fixed H2/CO flow without
disturbing the approximate mass percentages of 4, 9, and
18). After the results of the first series were analyzed,
the second series were run repeating the 60 SCFH H2/CO flow
rate and adding a test point at approximately 100 SCFH (30
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percent simulated dissociation). These two test points (18
percent and 30 percent dissociation) were run at both the
29.5 ft-lb condition and the 10.0 ft-lb test condition.
(Lower Hz/CO flow rates were used to maintain approximately
20 percent and 34 percent simulated dissociation conditions
at the lower load point.) Baseline testing was performed
before the first series of H2/CO testing, between the two
series, and after the second series. Standard fuel
consumption data was taken for all tests, however emissions
data were only taken on the first series of tests.
The results were analyzed by segmenting the equivalence ratio
into discrete regimes (see Table III-2). Within these small
regimes, data was treated as if all the data in that regime
had an equivalence ratio equal to the central value of that
discrete segment.* With this technique, data with only minor
differences in equivalence ratio could be more easily
evaluated with standard statistical routines. All of the
data reported (including data plots), then, represent the
average results within a given segment of equivalence
ratios. The only exception to this technique were the
emission results. Because of the limited amount of data, all
of the data is scatter-plotted on the plots of emission data.
Table III-2
Equivalence Ratio Segments
Segment Value
.455- .500 .478
.501- .545 .523
.546- .590 .568
.591- .625 .618
.626- .665 .647
.666- .735 .701
.736- .765 .750
.766- .805 .780
.806- .855 .830
.856- .904 .878
.905- .955 .940
.956-1.004 .999
*The equivalency ratio segments were selected based on
previous results with our specific A/F ratio control system
for the EFI system. The segments correspond to the specific
positions on the control box which accounts for the slightly
non-uniform spacing of the segments and equivalency values.
Most data tends to be near the central value of the segment.
See Appendix I for the derivation of the stoichiometric
fuel-air ratio.
-------
-19-
At the 29.5 ft.-lb. test point, even small amounts of
simulated dissociation gas show some change in engine BTE
performance. Figure 9 shows the results of 4 percent and 9
percent simulated dissociation. The lean limit was extended
somewhat with the HZ/CO gas, but no improvement in thermal
efficiency was observed at the 4 percent level. Only a
slight improvement was observed at 9 percent. Modest
improvements were observed at the 18 percent and 30 percent
level (see Figure 3).
B. Efficiency, Load 2
The 10 ft-lb load point results seem to have more mixed
results than the 29.5 ft-lb data. Table III-3 indicates
these results. The efficiency improvements with 20 percent
dissociation are what might be expected, but the negative
improvement in BTE with the 34 percent dissociation was
completely unexpected. If we arbitrarily group the similar
dissociation rates at the 10 and 29.5 ft-lb points (i.e. 18
and 20 become 20, and 30 and 34 become 30), we can compare
changes in efficiency between the two load points (see Table
III-4). It is apparent in Table III-4 that the improvements
with 20 percent dissociation at the 10 ft-lb load point are
similar to those at the 29.5 ft-lb point. However, if we
compare the 20 percent curve shape of the 10 ft-lb results
(Figure 10) to the shape of the 29.5 ft-lb data (18 percent
results - Figure 3), they bear no resemblance to each other,
and have effective slopes that are in opposite directions.
In fact, the shape and slope of the 34 percent dissociation
curve (10 ft-lb) which had a negative improvement bears more
resemblance to the corresponding 29.5 ft-lb data than the two
20 percent curves (for the two load points) which showed
similar efficiency improvements. This is somewhat puzzling
and a hypothesis is suggested later on describing a concept
of minimum required dissociation that may explain the slope
changes. No satisfactory explanation has been found for the
negative improvement in the 34 percent data.
C. Emissions
Emission tests were only run on the first series of tests (4,
9, and 18 percent dissociation at 29.5 ft-lbs). Results of
engine-out emissions (i.e. no catalyst) are shown in Figures
11, 12, and 13 for OM*, CO, and NOx respectively. The OM
results in Figure 11 for the baseline appear to indicate a
*0rganic Matter (OM) values were determined by a standard LDV
FID system. The OM values calculated include the weight of
oxygen in the hydrocarbon density of the exhaust. The
density value used was 37.69 g/cu ft (vs. 16.33 g/cu ft for
gasoline HC). No correction for the FID response to methanol
was applied.
-------
BRAKE THERMAL EFFICIENCY M/PHRTIAL DISSOCIATION
30
BRAKE THEHHAL EFFICIENCY H/PARTIHL DISSOCIATION
29
28
27
26
25
2M
23
22
21
20
y N
/ * >^-~~fc^"-JC^
/>' /^ ~* x-
' /
/
' S
1500 RPM
29.5 ft-lb
i i i i i i i i i i i
X BASEL INE INERT
HETHANOLI
N>7
*\1 H2/CO BTHRSS
N-3
T 9Z H2/CO BTHASS
N-3
S STOCK ICRSOL1NE.
9:1 CRI
N-l
O.MS O.SS 0.65 0.7S 0.8S 0.95 1 . OS
0.50 0.60 0.70 0.80 0.90 1.00
PHI
FIG 9
I 7
16
IS
I 1
13
12
I I
10
1500 RPM
10 ft-lb
BASELINE(NEAT
HETHANOLI
N-6
18* H2/CO BT HASS
N'S
0 30Z H2/CO BT HRSS
N-7
I
NJ
O
O.MS O.SS 0.6S 0.7S O.BS 0.9S I . OS
0.50 0.60 0.70 0.80 0.90 1.00
PHI
FIGURE 10
-------
-21-
Table III-3
Simulated Dissociation at
1500 RPM and 10.0 £t-lb
% Dissociation
Energy basis
Mass basis
Equivalence Ratio
(for best results)
Best Efficiency
Thermal
Efficiency (% BTE)
System
Efficiency (% SYE)
Difference
% Improvement*
BTE
SYE
"Compared to neat methanol
Simulated Dissociation
Weight Fractions
Baseline 20% 34
22.48 37.
19.98 33.
.750 .647
14.05 14.85 13.
14.05 15.35 14.
0 .50
5.7% -2.
9.3% 3.
baseline
%
41
98
618
72
51
79
3%
3%
Table III-4
Comparison of Efficiency Changes* at
Two Load Points with Simulated Dissociation
Mass Fraction
Dissociated 10 ft-lb 29.5 ft-lb
-BTE
20% +5.7% +5.3%
30% -2.3% +4.6%
-SYE
20% +9.3% +8.6%
30% +3.3% +10.1%
* Compared to neat methanol baseline,
-------
20
IS
10
OH EMISSIONS H/PAATIAL 0!5 SOCIRTI ON
BASELINE
1500 RPM
29.5 ft-lb
X BASELINE
N-3H
HZ HZ/CO BT
9X
N- I I
Q I8X
N-IB
MASS
H2/CO BT MASS
M2/CO BT HASS
10
9
8
7
c
z 6
a-
£ S
X.
JC
S M
S 3
2
CO EMISSIONS M/PARTIAl DISSOCIATION
BASELINE
1500 RPM
29o5 ft-lb
.MS .SO 55 .60 .65 .70 .75 .80 .85 .90
PHI
.MO .MS .50 .55
FIG 1 1
NOX EMISSIONS H/PARTIAL DISSOCIATION
1500 RPM
29.5 ft-lb
MS .SO .55 .60 .85 .70 .75 .80 .85 .90
X BASELINE
N*MX HZ/CO BT MASS
*9X M2/CO BT MASS
o"l8X M2/CO BT MASS
N-18
X BASELINE
N-JM
MX H2/CO BT MASS
N*8
9X HZ/CO BT MASS
N-l I
0 IBX MS/CO BT MASS
N*IB
.60 .65 .70 .75 .80 .85 .90
PMI
FIG 12
I
M
NJ
FIQ 13
-------
-23-
classic case of lean misfire as the equivalence ratio is
lowered. Adding small amounts of dissociation gas (4 or 9%)
improves the lean misfire limit slightly. Increasing the
amount of simulated dissociation to 18 percent, changes
substantially the lean misfire limit behavior of the engine.
The CO trends (Figure 12) seem to be similar to the OM
results. The baseline, 4 percent, and 9 percent dissociation
results are for practical purposes the same, while the 18
percent dissociation results are markably different. A
hypothesis is advanced later in this report that may explain
why the OM and CO results with 18 percent dissociation appear
to be different from the results observed for the baseline
and lower dissociation percentages.
The NOx results are somewhat less ordered than the OM and CO
data. The data in Figure 13 seems to indicate that at the
higher equivalence ratios there might be a slight difference
in the NOx production levels between using the simulated gas
versus using only neat methanol. However, there appears to
be no dependence of the NOx level on the amount of H2/CO
gas used. Further, the emissions levels with and without the
gas seem to coalesce at the leaner equivalence ratios. NOx
levels still appear to depend primarily only on the
equivalence ratio used. Since NOx formation can be related
to peak combustion temperatures, the data follows the
standard combustion theory, and highlights the NOx reduction
potential of operating at very lean equivalence ratios.
-------
-24-
VI. Discussion
A. Brake vs. System Efficiency
In this report, a distinction is made between Brake Thermal
Efficiency (BTE) and System Thermal Efficiency (SYE). Both
efficiencies are defined in this report as energy-out per
unit time divided by energy-in (E-in) per unit time. The
difference is in the manner of determining the value for the
amount of energy-in. For BTE, energy-in (per unit time) is
defined as the BTU/hr of the simulated dissociation gas con-
sumed plus the BTU/hr of methanol consumed (Equation 1).
(1) (E-in)BTE = (BTU/cu ft)(cu ft/hr)Mix + (BTU/hr)METH
For SYE a different approach was taken. It was assumed that
due to conservation of mass, one pound of methanol would
produce one pound of H2/CO dissociation gas. Therefore,
the density of the mix could be used to convert the volume
flow rate of the H2/CO simulation gas into a mass rate of
gas which would be equivalent to the same mass rate of
methanol. This effective mass rate of mix can then be added
to the mass rate of methanol used. Multiplying the total
mass rate consumed by the energy per unit mass of methanol
results in the energy-in (E-in) per unit time (Equation 2).
(2) (E-in)svE = C(cu ft/hr)(lb/cu ft)MIX + (lb/hr)METH] (BTU/lb)METH
The issue between the two methods for computing energy-in is
one of control volumes. In our case, we were simulating
dissociation with a mixture hydrogen and carbon monoxide
(Hz/CO). Therefore the control volume had two energy
inputs -- the H2/CO gas, and the methanol. The total
energy-in was the sum of the energy contained in the amount
of H2/CO gas consumed, plus the energy contained in the
amount of methanol consumed. The energy of the H2/CO gas
is calculated from the energy in the pure components, and was
proportioned by the volume percentage of the component.
These values are listed in Table III-l. For this situation
Equation 1 would be the appropriate approach to determine
energy-in for the purpose of determining what we have chosen
to define as brake thermal efficiency (BTE).
In the case of on-board dissociation, the system control
volume has only one energy source methanol. The creation
of the Hz/CO gas from the dissociation catalyst and heat
occurs within the control volume. Therefore, in this case it
is appropriate to consider Equation 2 as the proper method to
determine the amount of energy-in. By using equation two in
this manner, we can predict the engine performance as if we
had used an on-board dissociation reactor. Note however,
-------
-25-
that Equation 2 assumes that the conversion process from one
pound of methanol to one pound of dissociation gas is
complete. Therefore, Equation 2 represents an upper limit of
improvement which would be modified by the completeness of
the dissociation. In order to distinguish the actual engine
performance from the upper limit of predicted performance, we
defined the actual performance as Brake Thermal Efficiency
(BTE), and the upper limit as System Thermal Efficiency (SYE).
Creating this distinction between BTE and SYE leaves us in
the somewhat awkward position of having two energy values for
the same compound (see Table IV-1). The difference between
the two energy values is approximately 16 percent (for
complete dissociation). What this information suggests is
that the energy enhancement of dissociating methanol
completely is approximately 16 percent over that of running
the engine entirely on H2/CO gas supplied from some other
source.
Table IV-1
Effective Net Heating Values*
Units BTE SYE
w/simulated w/actual Diff.
mixture dissociation (BTE/SYE)
BTU/lb 10,039 8644 16.1%
BTU/cu.ft. 285 245.4 16.1%
*see Appendix III
The testing for this evaluation used the equivalent of only a
portion of the total engine fuel flow for dissociation. An
important question in this evaluation was then, "would small
amounts of dissociated fuel (assuming high dissociation
efficiencies) be beneficial?" If we assumed that
improvements in SYE over BTE were linear with the amount of
dissociation (assuming that the percent dissociation
represents complete dissociation of an equivalent amount of
liquid fuel as opposed to the partial dissociation of a
larger amount of fuel), we could predict the upper limit of
improvement of an onboard reactor over our simulation with
H2/CO gas. A comparison of this linear projection versus
the actual results in Table IV-2 suggest that the improvement
is not linear. The actual results from this limited data
average a 7.3 percent improvement over the linear
projection. Therefore, it appears that the change-in-state
energy change allows the engine efficiency to improve more
-------
-26-
rapidly with lower levels of dissociation. But because the
overall improvement due to the change-in-state is limited,
the rate of improvement [(SYE-BTE)XBTE] due to energy
enhancement will therefore decrease as the fraction of total
fuel that is dissociated approaches 100 percent. This
analysis, then, tends to support the intuitive conclusions
one draws from Figure 5.
Table IV-2
Effects of Energy Enhancement on
SYE Compared to BTE With H2 /CO
Linear Projection Actual*
Load 18% 20% 30% 34% 18%/20% 30%/34%
10 ft Ib - 3.2% - 5.5% 3.4% 5.8%
29.5 ft Ib 2.9% - 4.8% - 3.1% 5.3%
* From Table II-l and Table III-3, (SYE-BTE)XBTE.
Reviewing the theoretical increase in overall engine
efficiency due to dissociation, we find there are two
apparent sources of efficiency improvements in the overall
results. First there are the efficiency improvements
strictly from the addition of the H2/CO gas. This
improvement appears to be primarily due to the ability to run
the engine leaner as indicated in our BTE results (see Figure
3). Modest BTE improvements at these low H2/CO flow rates
are possible (around 5 percent improvement - see Tables II-l
and III-3). The second source of efficiency improvement is
from the increased energy potential of the fuel which would
be created by an on-board dissociation reactor (i.e., the
difference between BTE and SYE). Table IV-2 and Equation 2
suggest the degree of this second form of improvement is a
function of the amount of fuel dissociated and not
necessarily the equivalence ratio. Potential improvements
due to this second phenomenon at our low simulated rates are
of the same magnitude as the first effect (4 percent average
improvement - see Table IV-3). However, the relative
contribution of the second effect to the overall efficiency
would be expected to decrease as the fraction of dissociation
increases since this affect appears to be limited to
approximately 16 percent overall improvement (Table IV-1).
Also, the improvement would be an upper limit of improvement
for our dissociation rates since the analysis assumes
complete conversion efficiency for each fraction of total
engine fuel flow dissociated. The overall efficiency
-------
-27-
improvement from dissociating methanol is then the sum of the
two effects - hydrogen addition and the energy enhancement -
one a function of the fuel-air ratio that can be achieved
with the benefit of a given fraction of dissociation, the
other a function of the fraction dissociated.
BTE
SYE
Table IV-3
Sources of Efficiency Improvement
Source
(hydrogen addition)
(simulated energy enhancement)
overall (simulated dissociation)
Average Improvement*
5.2
4.1
9.3
*Average values from Table III-2 (30 percent dissociation
data at 10 ft-lb was not included in the average)
B.
Lean Limit
In discussing engine operation, it was noted that it was
difficult to operate the engine at an equivalence ratio
leaner than 0.58 under any of the dissociation fractions
tested. We had not expected this result since the engine was
baselined at as low as 0.62 equivalence ratio on neat
methanol (see Figure 2), albeit at a lower efficiency than
with the dissociation gas. Prior to testing the simulated
dissociation, we expected that there might be sufficient
hydrogen to extend the lean limit more than the extension
observed.
A review of explosive limits indicates that the lower limit
for methanol is around 6 percent by volume (9). This would
translate into an equivalence ratio of 0.46 (see Appendix II)
for the limits of combustion. From this information and
information from a contract with Ricardo (12) (which
demonstrated good efficiency with an HRCC methanol engine at
an equivalence ratio down to 0.6), we expected that adding
hydrogen (which has a lean limit of 0.1) to the engine would
improve the lean operating limit.
the operational limit was not extended lead us
paths, some more fruitful than others. One
approach was to evaluate the partial equivalence ratio of the
hydrogen, and the partial equivalence ratio of methanol. The
was to evaluate the relationship of each compound to
The fact that
down several
concept
-------
-28-
its flammability limit as the overall mixture was leaned.
However, the comparison of our data and that from the
relevant literature to the partial equivalence ratio approach
was inconclusive. In general, not enough data could be
gleaned from the literature to fully evaluate this
hypothesis, but the available information suggested that this
particular approach may not be worth pursuing further.
The exercise did, however, provide an important perspective
on the necessity to do all comparisons based on the weight
fraction of the fuel components. There is a reasonable
amount of literature on hydrogen supplemented engine
operation (3)(4)(5)(6)(7)(8). However, the majority of the
work used gasoline (generally Indolene) which was
supplemented by hydrogen or some mixture containing
hydrogen. In most cases, the data were only available in the
form of hydrogen energy fraction (HEF). In order to compare
these data to our data it was necessary to convert the HEF to
a hydrogen weight fraction (HWF). The following formula* was
used to convert HEF to HWF for mixtures of hydrogen and
gasoline or hydrogen and methanol.
(3) HWF = R (1/[(1/HEF) + R - 1])
where
HWF = hydrogen weight fraction
HEF = hydrogen energy fraction
R = the ratio of the energy content of the base fuel to
the energy content of hydrogen
Component Energy content (BTU/lb)**
H2 51743
Methanol 8644
Indolene 18579
By plotting this eguation (Figure 14), we can see an
interesting, but probably obvious, phenomenon. If we were to
present the our data in terms of hydrogen energy fraction
(HEF) as in Parks' (3) and MacDonald's (4) papers (MacDonald
also provided HWF), we would have significantly different
* Derivation is from HEF = HE/TE
where HE = hydrogen energy = (HWF)(energy content H2)
TE = total energy = (HWF)(energy content H2) +
(1-HWF)(energy content base fuel)
**see Appendix III
-------
-29-
H2 ENERGY FRflCTION VS. H2 HT. FRRCTION
0.0
0.0 0.2 0.4 0.6 0.8
H2 HT. FRPCTION (HHF)
FIG 14
T-VflRa I
INOOLINE + H2
N-21
T-VPR= MET
METHflNOL * H2
N-21
1 .0
-------
-30-
weight fractions for the same energy fraction because the
major component in our fuel differs from the fuel used by
Parks and MacDonald (methanol versus Indolene). Therefore,
because the stoichiometric fuel-air ratio must be determined
on a weight basis, it was necessary to make comparisons with
the literature references based on the hydrogen weight
fraction (HWF), and not the hydrogen energy fraction (HEF),
in order to compare the results when different fuels or mix
of fuels was used.
This insight lead to the possibility that we might be able to
correlate the engine lean limit with the HWF, and ignore the
effect of the CO on the lean limit. Table IV-4 tabulates
these data from references (1), (3), (4), and (5). (Note HEF
values from these references were transformed into HWF values
by the equation 3.) Plotting the data from Table IV-4 for
the mixture of hydrogen and methanol or gasoline (Figure 15) ,
we appear to have a reasonable correlation* between lean
limit and the hydrogen weight fraction (HWF) of the fuel. We
further note that extrapolation of these correlation lines
for either methanol or gasoline in Figure 15 (R2 = .9 to
.97) to an HWF of 1.0 (100 percent H2) would not yield the
lean limit value of 0.18 (3) for pure hydrogen. This suggests
that the improvement in lean limit is not linear over the
entire range of hydrogen supplementation and the
non-linearity is indicated in Figure 16. Even so, because of
the good correlation of the data in Figure 15, we can infer
that the linear regression is a reasonable approximation of
the behavior over the limited range of HWF's investigated.
We should also note that the equilibrium HWF which would
result from complete dissociation is limited to a value of
approximately 0.126 because of the presence of the carbon
monoxide.
From Figure 15, it is relatively easy to surmise why we were
having difficulty operating the engine leaner than 0.58
equivalence ratio. Even with 30 percent mass fraction of
Hz/CO simulated dissociation gas, we had only 3.6 percent
mass fraction of Hz (HWF) . At this level of HWF, Figure 15
seems to suggest that if we wish to significantly extend the
lean limit, then the HWF must be increased. If the HWF were
increased up to the maximum allowed by complete dissociation,
the equivalence ratio at the lean operational limit of the
engine might be expected to approach the flammability limit
of the dissociation mixture, but the degree would depend on
the engine design.
*Note: Because data from Houseman's paper did not appear to
agree with the other gasoline data (possibly because of the
effects of other constituents) it was not included in the
linear regression of gasoline-hydrogen mixtures. A separate
regression of Houseman's data by itself correlates well with
HWF.
-------
-31-
Table IV-4
Lean Limit Equivalence Ratio
vs .
Hydrogen Weight Fraction
Investigator /Fuel
Clemmens (Methanol)
Hirota (Methanol) (1)
Parks (Gasoline) (3)
MacDonald (Gasoline) (4)
Houseman (Gasoline) (5)
HWF
0
.005
.011
.022
.036
.112
0
.051
.097
.25
.144
0
.023
.046
.068
Lean Limit
.618
.618
.618
.583
.583
.435
.7
.6
.5
.34
.42
.71
.61
.59
.52
Parks* (H2) l.o .18
* Parks (3) indicates (in his Table B-l) that the lean limit
with pure hydrogen is 0.18 in a CFR engine.
-------
. 8
M2 HT. FRBCTION VS. IE1N L I "I
HI. FRBCTION VS. LEBN UNIT
APPROXIMATE LIMIT
METHANOL DISSOCIATION
T-VBR. LL1 NET
X HIX Of
METMBNOL.M2,BND CO
N-6 R50-.97
BO-.63 B1--I.93
T-«BH. LL1 GflS
INOOLINE H2.
flEf. 131 CHI
N-S KSO-.90
BO- 68 Bl -.77
.05 .10 .15 .20
12 M f. FPflCT ION
.25
. 30
.0
0. 0
0.2 0.4 0.6 0.8
H? MT. FHRCTION
T-VBR. LL2 MET
X MIX OF
METHBNOL.M2.BND CO
N>7
T-VBM* LL2 CBS
IMOOLINE n
REF. i3i i mi
N.6
T-VflR. LL2 MOUS
0 INOOLINt - M2.
REF. (SI
N.5
I . 0
KJ
I
FIG 15
FIG 1 6
-------
-33-
C. Minimum Required Dissociation (MRP)
In determining the fraction of dissociation required to
obtain meaningful results, previously we had seen in Figure 3
and Figure 9 that 18 percent and 30 percent simulated
dissociation provided somewhat greater improvements in BTE
over smaller fractions of dissociation (i.e., stock, 4
percent, and 9 percent). It is probably intuitively obvious
that the more H2/CO utilized in the engine (ie. higher
HWF) , the better the BTE would be. The data does tend to
support this conclusion, but when separating the data by
equivalence ratio groups versus the fraction of dissociated
methanol, an interesting picture emerges. The limited data
plotted in Figure 17 suggest that a minimum required
dissociation (MRD) rate of around 12.5 percent to 13 percent
apparently is necessary to show an improvement in BTE with
leaner mixtures. The same results appear in Figure 18 for
SYE which is on the same basis as Hirota's (1) data. Below
about 13 percent, and at extremely lean mixtures (less than
0.65), the projected BTE results (Figure 17) are
substantially lower than optimum neat methanol results (26.6
percent BTE, see Table II-l) . Above the apparent 13 percent
minimum dissociation value, the leaner mixtures are projected
to outperform neat methanol with the exception of the 0.7
equivalence ratio. It is interesting to note that for this
ratio (see Figure 17) the slope of the regression line of BTE
versus percent dissociated is nearly a horizontal line (0.01
BTE points per fraction of H2/CO). The slope of regression
line suggests that regardless of how much of the fuel
delivered to the engine is dissociated methanol, the thermal
efficiency remains relatively constant (at an equivalence
ratio of 0.7). When considering this situation from a system
efficiency basis (Figure 18), the outlook is somewhat more
promising with a projected 0.05 SYE point increase for each 1
percent increase in dissociated gas. If the linear
relationships shown in Figures 17 and 18 were valid for the
entire range up to dissociation of 100 percent of the total
energy supplied, we would expect a 3.6 percent improvement in
BTE (18.8 percent SYE) at 0.7 equivalence ratio. An 18.8
percent improvement in SYE over our base of 26.5 percent BTE
with neat methanol would result in an SYE of 31.61 percent.
This value is in close agreement* with Hirota's (1) results
* Values interpolated from Hirota's paper are 30 percent SYE
for 0.769 equivalence ratio, and 32 percent SYE for 0.667
equivalence ratio. Our calculated value is 31.61 percent SYE
for 0.7 equivalence ratio.
-------
BRAKE THERMAL EFFIC I ENCT IBTEI VS PERCENT DISSOCIATION
28
27
26
25
24
/
^
'
/
/ u -"
/ x *.----
/ * " '
' / ..-*"' _ _
> f- 7^-''
-""'""/
/ 1500 RPM
' 29.5 ft-11
/
i i i i i i
X PHI-. 583
N-2
PHI ... 61 8
N-4 RSQ-.99
80-24. 18 B1-.2I
Z PHI -.64 7
N-5 RSO-.60
80-25.89 B1-.06
- PHI -. 701
80^26^63 _8J*1JL1
STSTEH TMERNAL EFFICIENCT ISTEI VS PERCENT DISSOCIATION
24
<
/
31
) 5 10 15 20 25 30 35 33
PERCENT H2/CO BT HASS J2
FIG 17 31
30
29
28
UJ
£ 27
26
25
30
29
28
26
25
24
0
STSTEH THERHAL EFF 1 C I ENC T (S TE) VS PERCENT DISSOCIATION 21
,'* .--'''
1500 RPM
/ 29.5 ft-1
1 1 I 1 1 1
5 10 15 20 25 30 35
X PHI-. 583
N-2 22
PHI-. 618
N-4 RSO-.99 21
80-21. 17 B1-.26
Z PHI-. 647
N-S RSO-.89
80-25. 74 81-. 1 1
PHI*. 701
N-5 RSQ-.80
80-26^63 _BJ-. 05
)
'~^*~~*'
S^f^^^
y<
jr
f
x PHI-. 583 -.see
N-3
T PHI-. 647 -. 667
N-6
0 20 40 60 80 100
10 30 SO 70 90
PERCENT H2/CO BT HASS
FIG 1 9
3
PERCENT H2/C.O B( 1A3S
FIG 18
-------
-35-
for 100 percent dissociation*. However, if the above analogy
is applied to the leaner equivalence ratios, the results
predicted by the slope of SYE improvement are implausible.
Therefore, one must assume that the change in SYE with
increasing fractions of dissociation is not linear.
By plotting Hirota's values with our data (see Figure 19), it
seems that for lean equivalence ratios (less than 0.7) the
data appears to follow an exponential curve. Therefore if
the improvement in SYE decreases exponentially with
increasing fractions of dissociation, the case of a
relatively small overall improvement in SYE (i.e., a
relatively flat curve) might be approximated by a straight
line as appears to occur for an equivalence ratio of 0.7.
For leaner mixtures, the curve appears to have more "bow" to
it, and the straight line analogy (based on a linear
regression of the data on the front part of the curve) falls
apart.
The exponential shape might support and explain the data in
Figures 17 and 18 (minimum dissociation to show improvement)
in the following manner. If combustion at below the apparent
inflection point of 12 to 13 percent dissociated fuel were
dominated by methanol combustion reactions then one would
expect the BTE for these lower dissociation levels and at
equivalence ratios leaner than 0.7 to reasonably duplicate
the baseline case (i.e. no dissociation). A review of Figure
20 suggests that although the rate of fall-off in BTE at
these lean equivalence ratios with low fractions of
dissociation is not as rapid as with neat methanol, the
fall-off does somewhat parallel the neat methanol results.
Above this inflection point, it is hypothesized that there is
sufficient hydrogen, along with a lack of methanol at these
lean ratios, to support a hydrogen initiated flame front
within the combustion chamber. The assumed hydrogen flame
front would then set off the lean methanol mixture. In other
words, one could phrase this hypothesis as "moderate
fractions of dissociation serve as an ignition enhancement
system for extremely lean methanol operation."
The data at this one load point is, of course, not sufficient
to predict if such an inflectionpoint would occur at other
speed and load points at the same value of MRD or even occur
at all. Recognizing that even if 100 percent of the energy
*Note: Hirota (1) used a similar engine to the one used in
our study. He used a Nissan L20 engine and we used a Nissan
Z20 engine, the short block assemblies are essentially the
same with the major difference in engines being in the
cylinder head.
-------
-36-
passed through a dissociated reactor, 100 percent of the
methanol would probably not dissociate under the most optimum
conditions (and even less at part throttle). It may be
critically important that: 1) there could be a minimum
amount of hydrogen necessary to achieve a hydrogen initiated
flame front when dissociated methanol is used for fuel, and
2) there could be a switch between the hydrogen initiated
flame front and a methanol initiated flame front at different
dissociation levels with different optimum fuel-air ratio
requirements for each regime.
The previously described "minimum required dissociation"
(MRD) hypothesis might also provide a possible explanation
for the mixed results at the 10 ft-lb load point. If we
compare the 20 percent data in Figure 22 for 10 ft-lb, not to
the 18 percent data in Figure 21 for 29.5 ft-lb, but instead
to the 9 percent data in Figure 20 (for 29.5ft-lb), we see
similar curve shapes. Recognizing that the apparent
inflection point between having sufficient dissociation gas
for lean operation and not having sufficient quantity of gas
was around 13 percent for 29.5 ft-lb, the 9 percent and 18
percent flow rates for the 29.5 ft-lb load point fall on
either side of this apparent inflection point (Figure 20 and
21). Is it coincidence that they also have opposite slopes
of their efficiency versus equivalence ratio curves in the
same f/a ratio regime? If there is a cause and effect
relationship occurring, then it might be possible that the
opposite slope of the 20 percent and 34 percent curves at 10
ft-lb load point (Figure 22) may indicate that the inflection
point for MRD may have moved from 13 percent at 29.5 ft-lb to
between 20 percent and 34 percent for the 10 ft-lb load point.
VII. Postscript
In summing up, we can say that this program demonstrated that
the basic question, "Is it possible to achieve a significant
portion of the theoretical potential of full dissociation
with a less complex system?" appears to be able to be
answered in the affirmative assuming that the fractional
dissociation equipment is less complex than the full
dissociation system. Overall, up to 40 percent of the
potential efficiency improvement was achieved with only a 30
percent fraction of dissociation. However, the actual
improvement was, under the best conditions, only a modest 10
percent improvement. Future work in this area should be
cognizant of the simulation effects (BTE vs. SYE), the
apparent dependency of the lean limit on HWF, and the
potential impacts of the MRD level.
-------
BRAKE THERHAl EFFICIENCY M/PRRTIAl DISSOCIATION
in
29
28
27
26
25
2>4
23
22
21
20
_
f \
/ * ^"""^^--^C^
/' /^ * x-
J /
/
' S
1500 RPM
29.5 ft-lb
i i i § i i i i i i i
X BASELINE
HETHRNOLI
N-7
« HZ M2/CO
N-3
T 9X H2/CO
N-3
INERT
BTMHSS
BTMHS5
S STOCK IGRSOLINE.
9:1 CR1
N-l
O.US O.SS 0.65 0.75 0.8S 0.95 1.05
O.SO 0.60 0.70 0.80 0.90 1.00
PHI
FIGURE 20
BROKE THEHMflL EFFICIENCY M/PBflTlflL DISSOCIBTION
29
28
27
26
25
21
23
22
21
20
^
^§*.
^*^_
^^^^
^^^r**1^-^^
/ ^^~~* T~
/
/
/ ,
6
e
1500 RPM
29.5 ft-lb
i i i i i i i i i i i
X BRSEUNE INERT
METHANOLI
N-7
» 187. H2/CO BT HH5S
N-5
0 SOX. H2/CO BT MRSS
N-5
S STOCK IGRSOL INE. 9: 1
CHI
N-l
0.1*5 0.55 0.65 0.7S 0.85 0.9S 1.05
O.SO 0.60 0.70 0.80 0.90 1.00
PHI
FIGURE 21
BARKE THERNRL EFFICIENCY U/PRRTINL DISSOCIATION
17
16
IS
14
13
12
I I
10
9
8
O.MS 0.55 0.65 0.75 0.85 0.95 1.05
O.SO 0.60 0.70 0.80 0.90 1.00
PHI
FIGURE 22
1500 RPM
10 ft-lb
X BRSELINE INERT
NETHBNOLI'
N-6
182 H2/CO BT MRSS
N-S
0 30X H2/CO BT MASS
N.7
U)
^J
I
-------
-38-
References
(1) Toshio Hirota, "Study of the Methanol-Reformed Gas
Engine", Nissan Motor Co., Ltd., received 30th
September, 1980, by the journal of Japanese Society
of Automotive Engineers (JSAE ISSN 0389-4304).
(2) Charles Fayette Taylor, The Internal Combustion
Engine in Theory and Practice, Volume 1 and 2, The
MIT Press, MIT Cambridge, Massachusetts.
(3) F.B. Parks, SAE 760099, A Single-Cylinder Engine
Study of Hydrogen-Rich Fuels, 1976.
(4) J.S. MacDonald, SAE 76101, Evaluation of the Hydrogen
Supplemented Fuel Concept with an Experimental
Multicylinder Engine, 1976.
(5) J. Houseman and F.W. Hoehn, SAE 741169, A Two-Charge
Engine Concept: Hydrogen Enrichment, 1974.
(6) W. Jordan, SAE 790678, The Influence of Hydrogen
Addition to the Air-Fuel Mixture on Otto Engine
Combustion, 1979.
(7) F. Schafer, SAE 810776, An Investigation of Hydrogen
to Methanol on the Operation of an Unthrottled Otto
Engine.
(8) F.W. Hoehn and M.W. Dowdy, "Feasibility Demonstration
of a Road Vehicle Fueled with Hydrogen-Enriched
Gasoline," Paper 749105, presented at the 9th
Intersociety Energy Conversion Engineering
conference, San Francisco, California, August 1974.
(9) Merck Index, 9th Edition, 1976, Merck and Co. Inc.
(10) Aeronautical Vest-Pocket Handbook, Pratt & Whitney
Aircraft, Eleventh Edition, Eighteenth Printing, May
1966.
(11) Informal conversation with Amoco Oil Co.
(12) Ricardo Consulting Engineers, Optimum Engine for
Methanol Utilization, EPA Contract Number 68-03-1647,
Final Report, EPA 460/3-83-005, April 1983.
(13) Tokuichi Inagaki, Toshio Hirota, and Zene Ueno,
Combustion and Emission of Gaseous Fuel from Reformed
Methanol in Automotive Engines, presented at the
Alcohol Fuels Technology Third Internation Symposium,
Asilomar, California, May 28-31, 1979.
-------
-39-
References (cont'd)
(14) Edward F. Obert, Internal Combustion Engines and Air
Pollution, Intext Education Publishers, New York,
1973.
-------
-40-
Appendix I
Stoichiometric Fuel-Air Ratio
A. Simulated Dissociation with Methanol
The Stoichiometric fuel-air ratio of the different
mixtures can be determined by variations in the
chemical equation for complete combustion. The
standard form for methanol is
1) CH3OH + (1.5) 02 -> C02 + (2)H20
Note: for simplification, we will consider O2 as
air with a mixture of (1)02 + (3.73)N2 +
(0.4)Ar, and a molecular weight of 138.09.
The basic dissociation reaction is
Heat
(AI-2) CH3OH ^ CO + 2H2
Catalyst
From equation AI-2, we note that the mole ratio of
hydrogen to CO 2:1. Therefore, if we were to add
dissociated hydrogen and CO to the basic equation
(AI-1), the hydrogen and CO would need to reflect this
2:1 split.
3) (l-x)CH3OH + (x)(2)H2 + (x)CO + (y)02 ~ (a)C02 + (b)H20
where:
x = the mole fraction of the dissociated gas
added.
Balancing AI-3, we find that the value for "y" remains
the same at 1.5 as in equation AI-1. Therefore, we can
say for an equilibrium balance of dissociated hydrogen
and CO, the Stoichiometric fuel air ratio is the same as
liquid methanol. However, in our case, our compressed
gas cylinders did not have the exact equilibrium split
of H2 and CO. Instead of the 2:1 H2 to CO split, we
had a 65.65%: 34.35% split of H2 to CO which works out
to be a 1.9112:1 split. Substituting this actual split
in to AI-3, we have
) (l-x)CH3OH+(x)(1.9112)H2+(x)CO+(y)02--(a)C02+(b)H20
-------
-41-
Balancing this equation, we find that "y" is a function
of "x" - the mole fraction of the dissociation gas.
(AI-5) y = 1.5 - (0.0444)(x)
Converting each reactant on the left side of AI-4,
substituting AI-5, and forming a ratio of the fuel
reactants to the air reactants in AI-4 provides us with
the stoichiometric fuel-air ratio for our split H2 and
CO with methanol
[(l-x)(32.043)+(1.9112) (x) ( 2 . 016) + (x) ( 28 . Oil) ]
(AI-9) (f/a)s =
[1.5 - (0.0444)(x)] (138.09)
Where the molecular weights are:
CH3OH = 32.043
H2 = 2.016
CO = 28.011
Air = 138.09
By substituting various values for the percent
dissociation (values for "x") in equation AI-9, the
effect of the percent dissociation on the stoichiometric
f/a ratio can be determined. For instance, for "x"
values of 9% dissociation (0.09) and 30% dissociation
(.30) the computed stoichiometric values are 0.1550 and
0.1558 respectively. Since these values compare to
within 1% of the stoichiometric ratio of 0.1547 for
liquid methanol, the slight offset in the stoichometric
ratio for the dissociated gas 'was ignored in this
report, and all equivalence ratios were based on the
liquid methanol value rounded off to 0.155.
The hydrogen weight fraction (HWF) for our mix of
hydrogen and CO from compressed cylinders plus liquid
methanol is determined by:
(AI-10) HWF =
[(l-x)(CH3OH)-t-(1.9112)(x)(H2)
where: the molecular weights from AI-9 would be
substituted for H2, CO, and CH3OH
For the case where the engine were supplied with 100%
gas from an equilibrium dissociation of liquid methanol,
the HWF would be found by:
-------
(AI-11)
-42-
HWF = (2)(H2)/[(2)(H2)
(CO)]
The HWFs in table form for the percent dissociations
used in this study with compressed H2/CO simulation
gas are in Table AI-1.
Table AI-1
Hydrogen Weight Fraction
% Dissociation
.04
.09
.18
.30
.34
1.0*
1.0**
HWF
.005
.011
.022
.036
.041
.121
.126
* Simulation gas in compressed gas cylinders
** Equilibrium dissociation of liquid methanol
B. Dissociated Methanol
Because dissociated methanol contains several
components other than Hz and CO, equation AI-9
is not appropriate to calculate the fuel-air ratio
of the products from a dissociation reactor.
However, Hirota's paper (1) provides a chemical
analysis of typical products from a dissociation
reactor. Using Hirota's volume analysis and a
modified form of equations AI-3 and AI-4, we can
derive the weight fraction of each component which
is necessary to balance a chemical equation with
these constituents. Table AI-2 lists Hirota's
analysis.
Table AI-2
Composition of Dissociation Gas (1)
Compound(i) MW V(i)* H(i) C(i
H2
CO
CH3OH
CO 2
CH3OCH3
CH4
2
28
32,
44,
46
016
Oil
043
Oil
070
16.043
,63
,24
,05
,04
,03
,01
2
0
4
0
6
4
0
1
1
1
1
1
0
1
1
2
1
0
*Volume fraction determined by analysis of dissociation gas (1)
-------
-43-
Rather than write out a long formula with these
compounds, we can write a simplified equation that just
counts the hydrogen, carbon, and oxygen in the fuel
irrespective of the chemical compound.
(e)C + (f)H + (g)0 + (y)02 = (a)C02 + (b)H2O
where :
e = 0.4 = the sum of the products of C(i) times
v(i)
f = 1.68 = the sum of the products of H(i) times
V(i)
g = 0.4 = the sum of the products of 0(i) times
Balancing equation AI-12, we have
a = 0.4
b = 0.84
y = 0.62
The stoichiometric fuel-air ratio is
the weight of the fuel reactants in
of the air reactants (note: 02 can
air with a MW of 138.09).
reactants is simply the number
reactant times the molecular
reactant (note: (g)0 (fuel
simply the ratio of
AI-12 to the weight
be considered to be
weight of the fuel
of atoms for each fuel
weight (MW) for that
oxygen) is considered
The
fuel). Mathematically this would be expressed as
(AI-13)
(AI-14)
(f/a)
[0.4(12.011) + 1.68(1.008) + 0.4(16)]
(0.62)(138.09)
(f/a)s = .1506
We note that stoichiometric value for Hirota's mixtures
requires approximately 2.6% less fuel than the liquid
methanol fuel-air ratio of 0.1547. We also note that
the hydrogen-carbon ratio of Hirota's data of 1.68H:0.4C
or 4.2:1 HC ratio is about 5 percent greater than
methanols 4:1 HC ratio. Therefore, we assume Hirota's
analysis has some round-off and approximation errors
associated with it. A high hydrogen content would also
drive the stoichiometric ratio lean from the liquid
methanol value. Because the offset of Hirota's
stoichiometric ratio from that of methanol is relatively
small, and there is the possibility that Hirota
corrected for the offset in reporting his results, we
chose to ignore the offset in the presentation of the
data.
-------
-44-
The hydrogen weight fraction (HWF) for Hirota's fuel mix
is determined by:
15) HWF = [H(i)][v(i)][H]/[(e)(C) + (f)(H) + (g)(0)]
Where:
-H(i) and V(i) are form Table AI-2 for H2
-The molecular weights for H, C, and 0 are
H = 1.008
O = 16.0
C = 12.011
-e, f, and g are from AI-12
(AI-16) HWF = .112
C. Gasoline
Stoichiometric fuel-air ratios for gasoline-hydro-
gen mixtures were calculated from the equation
listed in Appendix A of Parks' paper (3) - his
equation A-l. A ratio of the fuel reactants to
the air reactants was formed as indicated
Clo H18 + [24HEF/(1-HEF>] (H2)
(AI-17) (f/a) =
([14.5 + (12HEF/(1-HEF)]/PHI) (02+ 3.76N2)
For stoichiometry "PHI" is set equal to 1 and AI-17
becomes
138.25 + [24HEF/(1-HEF)] (1.008)
(AI-18) (f/as) =
14.5 + [12HEF/(1-HEF>] (138.09)
where: HEF = hydrogen energy fraction*
*When HEF was not given it was calculated from HWF
HEF = HWF (EH2)/[HWF(EH2) + (1-HWF)(EG)]
where:
EH2 = LHV of Hydrogen = 51743 (BTU/lb)
EG = LHV of Indolene = 18579 (BTU/lb)
-------
-45-
The stoichiometric fuel-air ratio calculated for the
various gasoline hydrogen mixtures listed in references
(3), (4), (5), (6), and (8) are listed in Table AI-3.
Table AI-3
Stoichiometric Fuel-Air Ratio
of Hydrogen-Gasoline Mixtures
Investigator
Parks (3)
MacDonald (4)
Houseman (5)
Jordan (6)
HEF
0
.13
.23
.48
.32
.06
.11
.16
.41
.74
1.00
HWF
0
.051
.097
.249
.144
.022
.044
.064
.20
.50
1.00
(f/a)
.069
.063
.0582
.0454
.0538
.0664
.0638
.0616
.0491
.0311
.0292
Hoehn (8)
.29
.13
.0551
-------
-46-
Appendix II
Equivalence Ratio at the
Lower Flammability Limit
The equivalence ratio at the flammability limit
was calculated for several compounds. Included
were methanol, H2 , CO, and Indolene.
Typically flammability/explosive limits are
presented on a volume fraction basis as opposed to
a weight fraction basis. Since the equivalence
ratio is based on a mass comparison, the volume
based flammability limits must be converted to a
mass based fuel-air ratio in order to obtain a
mass based equivalence ratio. The Merck Index (9)
lists the following volume based flammabilty or
explosive limits.
Table AII-1
Volume Based
Lower Flammability/Explosive Limits (9)
Compound Lower Limit (%)
Methanol 6%
H2 4%
CO 12%
Indolene (Gasoline) 1.3%
An assumption is made that these limits do not
change with pressure, or if they do change the
effect is relatively minor. Using this assumption
we can then convert the fuel volume fraction to a
mass fraction at STP conditions (T = 518.7°R and P
= 2116 Ib/sqft). First the weight density (w) is
computed with equation AII-1. The molecular
weight (MW) of the fuel is found in Table AII-2.
(AII-1) w = p/RT
where
(AII-2) R = 1544/MW
-------
-47-
Table AII-2
Molecular Weights
Compound
Methanol
H2
CO
Indolene
air
MW
32.043
2.016
28.011
106*
28.966**
(AII-3)
(AII-4)
(AII-5
Next we form an equation that describes the
components in 1 cubic foot of a fuel(f) plus
air(a) mixture at the volumetric based lower
flammability limit (LFL).
[(LFL)(cuft)(w)(lb/cuft)]f + [(l-LFL)(cuft)(w)(lb/cuft)]
= Ib mix/cuft mix
Since the fuel-air ratio is simply a ratio the
respective weights in AII-3, we can determine the
fuel-air ratio at the lower flammability limit by
(f/a) = [(LFL) (w)]f / [(1-LFL) (w)]a
Dividing equation AII-4 by the stoichiometric
fuel-air ratio of the fuel in question provides us
with the equivalence ratio (phi) at the lean
limit. The results are listed in Table AII-3.
phi
L M L
(f/a)LFL / (f/a)
Compound
methanol
Hz
CO
Indolene
(Gasoline)
Table AII-3
Equivalence Ratio at the
Lower Flammability Limit
STOICH (f/a)
.155
.0292
.4057
.069
LFL (vol %) (9)
6%
4%
12%
1.3%
phi@ LFL
.455
.100
.325
.700
*Reference (11), note the average molecule is C7.64 HI 4.13
**Reference (10)
-------
-48-
Appendix III
Energy Values and Derivations
Lower heating values (LHV) or net heating values used in this
report for pure compounds were obtained from the Merck Index
(9) and Obert (14). Values are listed in table AIII-1. The
methanol used was industrial grade methanol analyzed at
better than 99.9 percent methanol. The analysis for LHV of
the methanol actually indicated a value approximately 1.4
percent less than Obert. However, since this test program
involved comparative testing with the same fuel, the absolute
value of the LHV for methanol was not as important.
Furthermore, the magnitude of the difference was relatively
small and considered inconsequential relative to engine
variability. Therefore, we elected to go with Obert's value
which results in more conservative values for engine
efficiency.
The LHV for the simulated dissociation mix (H2/CO) was
determined by rationing and summing the ratios of the heating
values based on their weight fraction of the total mix. The
mass densities were computed from the perfect gas law as
indicated in equation AII-1. The weight fractions were then
computed as in equations AI-3 and -4. The weight fractions
listed with these equations are the values used to compute
the LHV ratios. The resultant LHV values of the mix are
listed in Table AIII-2.
Table AIII-1
Net Heating Values
of Pure Compounds
Compound Net Value*(LHV)
Hz (9) 51,743
CO (9) 4,328
Methanol (14) 8,644
Methanol (ASTM D-240) 8,528
Indolene (ASTM D-240) 18,579
*BTU/lbn
-------
-49-
Table AIII-2
Net Heating Value
of the Simulation Gas
H
2 CO Mixture*
Heating Value (LHV)
(BTU/lb) 51,743. 4328 10,039
(BTU/cu. ft.) 269 315 285
Density (Ib/cu. ft.) .0052 .0727 .02839
*Gravimetric analysis of the gas cylinders used to provide
the simulation gas indicated the cylinders contained 65.65
volume percent H2 and 34.35 volume percent CO. (Analysis
accuracy was +1 percent of point.) Heating value of the
mixture in BTU/cubic foot was determined by proportioning the
heating value of the pure components based on their volume
percentage of the mix. The same was true of the density
calculations.
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