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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                             -3-

             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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                             -4-

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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                             -6-

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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                              -8-
  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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                              -9-

(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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                             -10-

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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                             -11-

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

-------
                             -13-

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.

-------
                             -14-

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.

-------
        -15-
 Supporting Analyses



         and



Technical Information

-------
                             -16-

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.

-------
                             -17-

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

-------
                             -18-

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
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-------
                             -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.

-------
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-------
                             -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  inflection•point 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
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^
^§*.
^*^_
^^^^
^^^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

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                             -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.

-------