EPA-AA-TSS-85-
                 Technical Report
Cold Starting Spark Ignition Engines with Methanol:
           An  Analysisof  Current  Options,
          and Their Impact on Air Quality
                         By

                William B. Clemmens


                    August 1985
                       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

-------
                              -2-

I.   Introduction

II.  Executive Summary
                      Technical Analysis *******
                                                       * * * *
III. Unassisted Vaporization
       A. Introduction
       B. Required Properties
       C. Summary

IV.  Physical Vaporization (Heat Addition)
       A. Introduction
       B. Required Properties
       C. Heat Addition Options
       D. Summary

V.   Chemical Vaporization (Fuel Additives)
       A. Introduction
       B. Required Properties
       C. Summary

VI.  Starting Fuels
       A. Auxiliary Fuels
       B. Fuels made from Methanol

VII. Effect of Cold Starting Approach on Air Quality and
       the Certification Process

-------
                             -3-




I.    Introduction

-------
                             -4-




II.   Executive Summary

-------
                             -5-

III.  Unassisted Vaporization

     A.   Introduction

     In order  to start  a  Spark  Ignited (SI)  engine,  it  is
generally accepted  that  a  sufficient quantity  of fuel  must
vaporize to  provide  a combustible mixture  of fuel vapor  and
air  in  the  vicinity  of  the  spark plug  during the  ignition
event  (1).   Under cold  starting  conditions,  the  latent  heat
of  vaporization  and  the  curve  of   vapor  pressure  versus
saturation  temperature  play  important  roles  in  determining
the quantity of  vapor available  at the spark plug.   Gasoline
is a  multi-component mixture, and certain  components  in  the
fuel  can  provide sufficient  vapor  phase  fuel  under  cold
starting conditions,  provided enough  gasoline  is added to the
air charge.   Neat methanol,  of course, has  no  such volatility
additives,    and   has  a   fixed   volatility   curve.   Although
volatility  components  can  be  added  to  methanol,   in  this
section  we  want  to   evaluate   the   requirements  for  cold
starting neat  methanol  without  additives  or  other  starting
devices.

     In  this  situation,  essentially all  the  heat  used  to
vaporize the neat methanol  must  come  from the latent  heat  of
the  air  and  liquid  fuel  charge,  from heat   transfer  from
warmer  engine  surfaces   (if any), and  from  the work  of  the
compression  stroke;   most  will  come   from   the   work   of
compression.  The process can be  envisioned as  consisting  of
five steps:  1)  Liquid  fuel and  air  at the prevailing ambient
temperature  are  mixed.   2)  The  fuel   begins  to vaporize  and
the  temperature  of   the  air  and  remaining  liquid  fuel  is
depressed until  an  equilibrium is reached.   The  amount  that
can be  vaporized depends on the  initial  ambient temperature.
3) The  depressed mixture temperature  leads to  heat  transfer
from  engine  surfaces  which  have remained  at  the  ambient
temperature.  4)  The mixture is  then compressed  and heated,
with  some  heat  loss  to the cylinder  walls.   5) The  heat  of
compression  is  used  to   vaporize  more of  the methanol  until
equilibrium  is  reached.    In  actuality,  the  second,  third,
fourth, and fifth steps  overlap in time.

     B.  Required Properties

     Several   investigators   (2)(3)   have    used    various
approaches to model  the  vaporization  or  starting capabilities
Of neat  methanol.  Some  are more sophisticated than others.
Our  approach while  somewhat  simpler   than  others provides  a
different  perspective to  the issue   of  cold   starting  with
methanol.

     Our  approach was  based on  the   premise  that a  certain
quantity of vapor is  required for  the  starting  of  SI  engines,

-------
                             -6-

and  that  the  macroscopic  equilibrium conditions  to  maintain
this vapor must  at  least be met.   Otherwise,  some the  vapor
would  condense  and  reduce  the  vapor . fuel-air  ratio  to  a
non-combustible  mixture.   The  temperature  of   the   mixture
after  adiabatic  ideal gas compression  and  evaporation serves
as an  indication  of equilibrium reserve,  and as such  must  be
at or  above  the  saturation  or dew temperature  for the  vapor
phase  fuel  concentration.   Any  temperature  above  the  dew
temperature  would  be   a   temperature   reserve   (TR).   This
reserve would be utilized for  such things as heat  transfer  to
the  cylinder  walls,  non-ideal process  losses,  rate  effects
(i.e., insufficient time to  reach  equilibrium),  etc.   Because
these  losses  are always present and  must be  compensated,  the
temperature  reserve would  have to be  a positive number  in
order  to maintain the fuel  in  a vapor state under  equilibrium
conditions.   In  our  evaluation,  we  compared  the  calculated
temperature  reserve  (TR)  at  known  start capability conditions
to the reserve at unknown cold starting conditions.   As long
as  the  reserve   at  the  lower temperature  was  equal  to  or
greater than  the  reserve at the higher temperature (with  the
known  start  capability),  we   assumed   the  potential   for  a
successful cold start existed.

     This  approach  however  assumes  all  of  the  evaporation
necessary  to  achieve a  combustible mixture  occurs during  the
compression  stroke.   Before accepting  the  reasonableness  of
this  approach  (i.e.,  no  fuel  evaporation  in  the   intake
manifold or  in  the cylinder  during  the  intake  stroke when
cranking),   we  studied  the relationship  between  the  vapor
equivalence  ratio and the  ambient  temperature which  would  be
required to  support  a  given  vapor  ratio  after evaporation.
Because  methanol   is  a  single  component  mixture,   a  well
defined  relationship  exists between  the partial  pressure  of
methanol and  the  saturation  or dew temperature  for  methanol.
The  partial  pressure of a  given  fuel mixture  is  governed  by
the  concentration of the vaporized .methanol  (i.e. the  vapor
f/a  ratio),  and  the  local  pressure.   The  following  equation
from Appendix  I  in  reference  (5)  was  used  to  determine  the
partial  pressure of  methanol for  a  range  of  pre-selected
vapor equivalence ratios.

(III-l)          PV = (MAP)  (1/[1 + (7.155/VPHI)])

     where:
             PV   = partial p-ressure
             MAP  = manifold air pressure
             VPHI = vapor equivalence ratio

     The   saturation   temperature   is,    of   course,    the
temperature  that  separates  the  liquid phase   from  the  gas
phase, and in our  case  it is  the  temperature which  must  be

-------
                             -7-

maintained  in  order   to   prevent  our   pre-selected   vapor
equivalence ratio  from condensing.   Reference  (6)  describes
the  relationship  between   the   partial  pressure   and   dew
temperature as:

(III-2)          TDEW = (-1961.8678)/[log,0(PV) - 8.639821)]

     where:
             TDEW =(°K) = saturation temperature
               PV = (mm Hg) = partial pressure from Eq.  III-l

     Equation  III-2,  does  not consider  the  fact that  in the
case of an intake manifold, in order to  achieve  a  given vapor
equivalence ratio, an  equal  amount of  liquid methanol must be
evaporated.  The process of evaporating  the  liquid  fuel would
depress  the  surrounding ambient  temperature.   Therefore,  in
order   to   maintain   our   computed   TDEW   value    for   our
pre-selected vapor PHI  in  equation III-l,  our initial ambient
temperature before  evaporation must be  above the TDEW value
at  least  by  the  amount of  temperature depression  caused by
the  evaporation  of  the fuel.   The  following  equation  from
reference  (4)  was selected to determine the temperature drop
from the evaporating fuel.

(III-3)  TDROP =  [(x)(F)(HLG) + (Q)]/[(1-F+xF)(CP)](5/9)

     where:
            TDROP = (°C) = temperature drop
                x = portion evaporated,  in our case = 1.0
                F = fuel-air ratio
              HLG = heat of vaporization = 474 BTU/lb (ref. 4)
               Cp = is for a mixture of  fuel-air vapor,
                    reference  (4)   lists Cp   as  0.245  for  a
                    stoichiometric  mixture of methanol vapor
                    and  air,  and  a  Cp  of  0.240  for  air
                    alone.
                Q = heat addition, in our case = 0.0
                  - Heating or  cooling  of excess  liquid fuel
                    was neglected
                  - Pressure  changes  (if  any)  from cooling
                    were ignored

     By  substituting  the same pre-selected  equivalence ratio
that was  used in equation  III-l  into equation  III-3,  we can
identify the  temperature drop that  would  be  associated  with
vaporizing  a  given  amount  of  fuel  corresponding to the dew
temperature  calculated  in  equation  III-2.   Combining  the
results  of  III-2  and  III-3,  we  have  the  initial ambient
temperature before  vaporization  (which  is assumed  to  be the
same  as the  initial  intake  manifold   air  temperature)  that
would support the pre-selected vapor equivalence ratio.

-------
                             -8-

(III-4)           TAMB = TDEW +  TDROP

     Where:
             TAMB  = ambient temperature.
             TDEW  = saturation temperature,  equation III-2.
             TDROP = temperature drop,  equation III-3.

     Results  of   equation  III-4   for   a  range   of   vapor
equivalence  ratios  and  manifold  air  pressures  (MAP),  are
plotted  in  Figure  III-l  and  listed  in  Table  III-l  through
III-4.  The range  of  manifold  pressures  used to  derive  these
tables covered  the range  (suitable  for modeling purposes)  of
observed manifold pressure values for  a typical  four cylinder
engine  under  cranking  conditions   at  99kPa(wet)  barometric
pressure and a  range  of cranking speeds  from  200 RPM to  600
RPM.  From  Figure  III-l  and  the  Tables,  the   influence  of
varying  the  manifold  air  pressure  slightly from the  normal
cranking condition  (85 kPa) appears to have little  effect  on
the  amount  of  fuel  that  can  be   vaporized  in  the  intake
manifold  at  temperatures  below   0°C.   Large  changes   in
manifold pressure may have an  observable  affect  on the amount
of  fuel  that  can  be  vaporized,  but we will investigate that
prospect later.

     For the time  being,  these equilibrium  data  suggest that
if  we are  to  consider starting  a neat methanol  engine  at
ambient temperatures of -18°C  (0° F) or below  (TAMB  in Tables
III-l to  -4),  it  seems  safe to say that there will  be very
little evaporation occurring within  the  inlet  manifold during
cranking,  at  most  maybe  a 0.05  to  0.06  vapor  equivalence
ratio, and  at  -35°C,  maybe  a  0.03  vapor equivalence  ratio.
(Lack  of  time  may  prevent  even   these points  from  being
reached).   For   a   carbureted  engine,   the  lack   of   fuel
vaporization  in   the   inlet  manifold  could  cause  serious
distribution problems,  but  more  important for our  study,  the
lack  of  evaporation  in  the  inlet  manifold  means  that  the
majority of the  evaporation  to produce an ignitable  vapor  at
cold  conditions  must  occur during the   compression  stroke.
This  evaluation,   therefore,    suggests   that   our   initial
approach (i.e.,   no  fuel evaporation in the  intake manifold or
cylinder  during  the   intake  stroke  when   cranking)   is   a
reasonable assumption to begin our analysis.

     We have  so far  neglected the  fact  that  the temperature
drop caused by evaporation will  create a  temperature gradient
and  encourage   heat  transfer  from   the  manifold  or  cylinder
walls, leading  in turn  to  a higher vapor equivalence ratio.
Evaporation of enough fuel to cause a vapor equivalence ratio

-------
   VflPOR PHI VS  MRNFOLO  BIR  TEMP
1.0

0.9

0.8
uJ 0.6
o
z
l*J _ _
_j 0. 5
oc

30.4
UJ

g 0.3
a.
cc
» 0.2

  0.1

  0.0
   -60
                                               » VflP  PHI  H/EVflP
                                              RT  100.0 KPR
                                              =  VflP  PHI  H/EVflp
                                            RT 85.0  KPR
                                            N-29

                                              «  VflP  PHI  H/EVflP
                                            RT 75.0  KPR
                                            N-29
       -20    20    60    100   1UO   180
    -40     0    40    80   120   160   200
         RMBIENT TEMPERRTURE  (OEG C)
                          -  /

-------
       TABLE III-I


 VAPOR EQUIVALENCE RATIO
         TABLE III-2


   VAPOR EQUIVALENCE RATIO
MAP
(KPA)
100.
too
100
too
100.
100.
100.
too.
100.
100.
100.
100.
100.
too.
100.
100.
100.
100.
100.
100.
100.
100.
too.
100.
100.
too.
too.
too.
100.
.000
.000
.000
.000
.000
.000
.000
.000
.000
000
.000
.000
.000
000
.000
000
000
000
000
000
000
000
000
000
000
000
000
000
000
VAPOR
PHI
0.
0
0
0
0.
0
0.
0
o
0.
o.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
1 .
.005
.010
.020
.030
.040
.050
.060
.070
.080
.090
. 100
. 150
.200
.250
.300
350
400
450
500
550
600
650
.700
.750
.800
.850
.900
.950
.000
PART
PRESS
0
0
0
0
0
0
0
0
1
1
1
2
2
3
4
4.
5.
5.
6.
7.
7.
a.
8.
9
to.
10.
11 .
1 1 .
12.
.070
. 140
.279
.418
.556
.694
.832
.969
. 106
.242
.378
.053
.719
.376
.024
.664
.295
917
.532
. 138
.737
.328
.912
.488
.057
.618
. 173
.721
.262
TOEW
(OEG C)
-53
-45
-37
-32
-28
-25
-23
-20
-19.
-17
-15
-9
-5.
-2
0.
3.
5.
7.
9.
10.
12
13
14
15
16
18
18.
19
20
.221
.547
.328
.246
.510
.536
.056
.922
.047
.372
.858
.880
.499
.023
.865
.339
.503
.427
. 160
.735
. 180
.513
.750
.904
.986
.002
.961
.868
.729
TOROP
(DEC C)
0.
1 ,
3.
5
6.
8.
10.
1 1 .
13.
15.
17.
25.
34.
42.
51 .
59.
68.
76.
85.
93.
102.
110.
1 19.
127.
136.
144.
153.
161 .
170.
.850
.701
.401
. 102
.803
.503
.204
.905
.606
.306
.007
.510
.014
.517
.021
524
028
531
035
538
042
545
048
552
055
559
062
.566
.069
TAMB
(OEG C)
-52.
-43.
-33
-27.
-21
-17
-12.
-9.
-5.
-2.
1 .
15.
28.
40.
51.
62.
73.
83.
94.
104.
114.
124.
133.
143.
153.
162
172.
181
190
.371
.847
.926
. 143
.707
.033
.852
.018
.442
.066
. 149
.630
.515
.494
.886
.863
.531
.958
. 194
.273
.221
.058
.799
.456
.041
.561
.023
.434
.798
      TABLE I I 1-3

VAPOR EQUIVALENCE RATIO
MAP
(KPA)
75
75
75
75
75
75
75
75.
75.
75
75.
75.
75.
75
75
75
75
75
75
75
75.
75.
75.
75.
75.
75.
75.
75.
75.
.000
.000
.000
.000
.000
.000
.000
.000
.000
.000
.000
.000
.000
.000
.000
.000
.000
.000
.000
.000
.000
.000
000
000
.000
000
000
.000
000
VAPOR
PHI
0
0
0
0
0
0
0
0.
0.
0
0.
0.
0.
0.
0.
0
0.
0.
0
0.
0.
0.
0
0
o
0.
0
0
1
.005
.010
.020
.030
.040
.050
.060
.070
.080
.090
. 100
. 150
.200
.250
.300
.350
.400
.450
.500
.550
.600
.650
.700
.750
. BOO
.850
.900
.950
.000
PART
PRESS
0.
0
0
0
0
o
0
0.
0
0
1 .
\
2.
2.
3.
3.
3.
4
4
5
5
6
6
7
7
7
8
8
9
.052
. 105
.209
.313
.417
.520
.624
.727
. 829
.932
.034
.540
.039
.532
.018
.498
.971
.438
.899
.354
. 803
.246
.684
. 1 16
.542
.964
.380
.791
. 197
TDEW
(DEG C)
-56.
-48
-40
-35
-32.
-29
-26
-24.
-23.
-21
-20.
-14.
-9.
-6.
-3.
-1 ,
0
2
4
5
7
a
9
10
1 1
12
13
14
15
.259
.800
.817
.886
.263
.381
.977
.910
.094
.472
.006
.222
.985
.625
.835
.446
.644
.501
. 174
.694
.087
.373
.566
.679
.722
.702
.626
.500
.330
TDROP
(DEG C)
0.
1 ,
3.
5.
6
8
10
1 1 .
13.
15.
17.
25.
34.
42.
51 .
59.
68.
76
85
93
102
1 10
1 19
127
136
144
153
161
170
.850
.701
.401
. 102
.803
.503
.204
.905
.606
. 3O6
.007
.510
.014
.517
.021
.524
.028
.531
.035
.538
.042
.545
.048
.552
.055
.559
.062
.566
.069
TAMB
(DEG C)
-55.
-47.
-37.
-30.
-25.
-20.
-16.
-13.
-9.
-6.
-2.
1 1 .
24.
35.
47.
58.
68.
79.
89.
99
109
1 18
128
138
147
157
166
176
IBS
.409
099
.416
.784
.461
.877
.773
.005
.489
. 166
.999
.288
.029
.892
. 186
.078
.671
.033
.208
.232
. 129
.918
.615
.231
.777
.260
.688
.066
.399
MAP
(KPA)
85
85
85
85
85
85
85
85
85
85
as
85
85
as
as
as.
85.
85.
85.
85.
as.
85.
as.
85.
85.
85.
85.
85.
85.
.000
.000
.000
.000
.000
.000
.000
.000
.000
.000
.000
.000
.000
.000
.000
.000
000
000
000
000
000
000
000
000
000
000
ooo
000
000
VAPOR
PHI
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
1 .
.005
.010
.020
.030
.040
.050
.060
.070
.080
.090
. 100
.150
.200
.250
.300
.350
.400
.450
.500
.550
.600
. 65O
.700
.750
BOO
.850
900
.950
.000
PART
PRESS
0
0
0
0
0
0.
0.
0.
0.
1 .
1 ,
1 .
2.
2
3.
3.
4.
5.
5.
6.
6.
7.
7.
8.
a.
9.
9.
9.
10.
.059
. 1 19
.237
.355
.473
.590
.707
.824
.940
.056
. 172
.745
.311
.870
.421
964
500
030
552
067
576
079
575
065
548
026
497
963
423
TDEW
(DEC C)
-54
-47
-39
-34.
-30
-27.
-25
-23.
-21 .
-19.
-18.
-12.
-8.
-4
-1 .
0.
2.
4.
6.
7.
9.
10.
1 1 .
12.
13.
14.
15.
16.
17.
.948
.396
.312
.316
.645
.723
.286
.191
.349
.705
.217
.351
.052
.642
.810
615
737
623
321
865
280
586
799
929
988
984
923
811
654.
TDROP
(OEG C)
0
1
3
5
6
8
10
1 1
13
15
17
25
34
42
51
59
68
76
85.
93
102
1 10
119.
127.
136
144.
153.
161 .
170.
.850
.701
.401
. 102
.803
.503
.204
.905
.606
.306
.007
.510
.014
.517
.021
.524
.028
.531
.035
.538
.042
.545
.048
.552
.055
.559
.062
.566
.069
TAMB
(DEG C)
-54
-45
-35
-29
-23
-19
-15
-1 1
-7
-4
-1
13
25
37
49
60
70
81
91 .
101.
Ill
121.
130.
140.
150.
159.
168.
178.
187.
.097
.695
.910
.214
.842
.219
.082
.286
.744
.399
.211
. 160
.962
.875
.21 1
. 139
.764
. 154
.356
.403
.322
. 131
.847
.481
.043
.542
.985
.377
.723
      TABLE III-4


VAPOR EQUIVALENCE RATIO
MAP
(KPA)
65.000
65.000
65. OOO
65.000
65.000
65.000
6S.OOO
65.000
65.000
65. OOO
65.000
65.000
65.000
65.000
65.000
65.000
65.000
65.000
65.000
65.000
65.000
65.000
65.000
65.000
65.000
65.000
65.000
65.000
65.000
VAPOR
PHI
0.005
O.010
0.020
0.03O
O.O40
0.050
0.060
0.070
0.080
0.090
O. 100
0. 150
0.200
0.250
0.300
0.350
0.400
0.45O
O.SOO
0.550
0.600
0.650
0.700
0.750
0.800
O.B50
O.900
0.950
1 .OOO
PART
PRESS
0.045
O.091
0. 181
O.271
0.361
O.451
0.541
0.630
0.719
0.807
O.896
1 .335
1 .768
2. 194
2.616
3.031
3.441
3.846
4.246
4.640
5.029
5.413
5.792
6. 167
6.537
6.902
7.263
7.619
7.971
.TOEW
(DEG C)
-57.739
-50.383
-42.515
-37.656
-34.088
-31 .249
-28.882
-26.847
-25.059
-23.463
-22.020
-16.328
-12. 161
-8.857
-6. 1 13
-3.765
-1.711
0. 115
1 .759
3.252
4.621
5.885
7.057
8. 150
9. 174
10. 136
1 1 .044
11 .903
12.717
TDROP
(DEG C)
0.850
1 .701
3.401
5. 102
6.803
8.503
10.204
1 1 .905
13.606
15.306
17.0O7
25.510
34.014
42.517
51 .021
59.524
68.028
76.531
85.035
93.538
102.042
1 10.545
1 19.048
127.552
136.055
144.559
153.062
161 .566
170.069
TAMB
(DEG C)
-56.889
-48.682
-39.113
-32.554
-27.285
-22.745
-18.678
-14.942
-1 1 .454
-8. 157
-5.013
9. 182
21.853
33.661
44.907
55.760
66.317
76.646
86.793
96.790
106.663
1 16.429
126. 105
135.702
145.229
154.695
164. 106
173.468
182. 7B6
                                                                                                                                                         1
                                                                                                                                                         o
                                                                                                                                                         I

-------
                             -11-

of 0.05 at -19°C and 85 kpa MAP  implies  a  mixture temperature
drop  of  about  8.5°C  which  seems  small  enough  that  the
resulting heat transfer can initially be ignored.

     If  there  is  essentially no  temperture  drop and  no heat
transfer, then from  an  equilibrium  point of  view, evaporation
before  compression  does not affect  the  energy balance  which
determines the  conditions  at the end of compression,  so such
evaporation can be  ignored  in  the   equilibrium  calculations.
With these  considerations  in mind,   in  order to  evaluate the
effect   on   equilibrium   conditions   of  the   incrementally
increasing  pressure   that  occurs   as  the  piston  travels
upwards, we began by  reevaluating equation III-2  and equation
III-l.  Equation III-2 can be written as
(III-5)      TDEW = A/tlogt 0(PV)-B]

     where:
             A = -1961.8678
             B = +8.6339821

     Performing  some  algebra  on  III-5,  we  can  write  the
solution for PV as:

(III-6)      PV = 10J

     where :
             J = (A/TDEW) + B

     If we  assume  for  this analysis that the initial pressure
levels in the cylinder under cranking  conditions  are the same
as  the  average  intake  manifold  conditions,   then  we  can
substitute  for PV  in the partial pressure equation (equation
III-l) with the relationship identified in III-6.

(III-7)      10J = (MAP) (1/[1 + (7.155/VPHI)] )

     Solving for VPHI, we have

(III-8)      VPHI = 7.155/[(MAP/10J) - 1]

     Solving III-8 by  substituting  selected  values  of TDEW in
the "J" identity  results in approximately the same  values  as
in Tables  III-l  through  III-3.   (round-off  errors are assumed
to be  the  cause  of  the  small  differences  occuring  in  the
third  decimal   place).   However,  equation  III-8,   is   now  a
general  equation  relating  pressure   (MAP)   and  saturation
temperature in  the "J"  term   to the vapor  equivalence  ratio
(VPHI).  Further,  if we solve  equation  III-4  for   TDEW,  and
substitute  these results in  the expression  for  "J", we have a
relationship that now  includes  the  effect of the temperature

-------
                             -12-
drop  from  the  vaporizing  fuel,  and  the  ambient
necessary to maintain the vapor equivalence ratio.
                                                   temperature
(III-9)

     where:

(111-10)
             J = [A/(TAMB-TDROP)1  + B
             TAMB = (°K)
             TDROP = (5/9)(1/Cp)[(f/a)s(VPHI)(HLG)]  = (°K)
             (f/a)s = 0.155 = stoichiometric fuel air ratio
             VPHI = vapor equivalence ratio
             HLG = 474 BTU/lb = heat of vaporization (ref. 4)

     The  relationship  for   "J"  can  further   be modified  to
include the  adiabatic temperature which  would reflect  a  new
saturation temperature as the cylinder pressure is increased.
(III-ll)
             J = [A/(TAD - TDROP)] + B
     where:
             TAD = adiabatic temperature (°K) =TAMB (p2/MAP)R
             TAMB = (°K) = ambient temperature
             p2 = new cylinder pressure
             R = (k-l)/k
             k = 1.38 = ratio of specific heats for a
                         methanol and air mixture (4)
             MAP =   manifold air  pressure which is  assumed
                     be  the  same  as  the  initial  conditions
                     the cylinder during cranking conditions.

     Putting everything together, we have

(111-12)         VPHI = 7.155/[(pz/10J) - 1]
                                                              to
                                                              in
     where
(111-13)
             pz = (mm Hg) = cylinder pressure (divide kPa by
                             0.133224 to obtain mm Hg)
             J = (A/[TAD - (5/9)(l/Cp)(f/a)s (VPHI)(HLG)]) +B
             A = -1961.8678
             B = +8.6339821
             CP = .240
             (f/a)s  = 0.155
             (HLG)  = 474 BTU/lb

             TAD  =  (°K) = TAMB (pz/MAP)R = adiabatic temp-
                     erature   resulting   from  increasing   the
                     cylinder  pressure  from  MAP  to  pz.   For
                     our  analysis,  p2  was- incremented  from  the
                     MAP value to determine TAD values.
             TAMB =  (°K) = ambient temperature
                R =  (k-l)/k
                k =  1.38

-------
                             -13-

     With equation  111-12,  we now have  an  equation*  that  allows
us  to   increase   the  cylinder   pressure   from  the   manifold
conditions  for a   given  ambient  temperature,   and  observe  the
resultant vapor equivalence ratio that would be  allowed  to  occur
under equilibrium  conditions.   If the  vapor  ratio goes down  as
the pressure  is increased,  then  we can assumed  that  some of  any
pre-vaporized  fuel  will condense,  if it goes  up we can  assume
that any  fuel  vaporized in the  intake  tract (in  this  analysis)
will remain vaporized.

     Figure  III-2  shows the  results  of solving  equation  111-12
for  four   ambient   temperature   levels.    From  85  kPa   (the
approximate  MAP  observed  during  cranking)  to over  3500  kPa
(absolute   pressure),   we   see   a   steadily   increasing   vapor
equivalence  ratio  that  is  supportable  under  the  equilibrium
conditions  given.   Equation  111-12  also allows us  to  explore
the concept of temperature reserve  (TR)  as it relates to  the
potential   for  predicting  the   upper   limits   of   cold   start
performance   with   neat  methanol.    The  assumption   is   that
equilibrium   values   will    define   the   maximum   permissable
performance,  and  other factors such  as  heat  transfer rates,  and
allowable fuel  evaporation  rates will  only serve to  reduce  the
maximum perfomance.   The magnitude  of these compromising factors
and their  influence is for  the  time being  ignored  in  order  to
determine the upper limit of performance.

     Since  it  is  generally  accepted that  a  certain amount  of
fuel vapor   is  necessary  to   initiate   combustion,  we  can  use
equation  111-12 to  determine  an  equilibrium temperature reserve
in  the  following  manner.    First,  we   initially   define  the
temperature   reserve    (TRX)   as   the  difference   between  the
temperature  that   theoretically   exists  in  the  cylinder  after
adiabatic compression  (TAD)  minus  fuel evaporation  temperature
losses   (TDROP)  and minus  the saturation  temperature (TDEW)  of
the  vapor  mixture,  without  accounting  for heat  losses to  the
cylinder walls.

(111-14)     TRX = (TAD - TDROP)  - TDEW

     Note,  that  the  adiabatic temperature  and the  temperature
drop are related  (in  the  "J"  term)  to  the  vapor  equivalence
ratio (VPHI)  for  a given  cylinder pressure by  equation 111-12.
The  calculated   vapor  equivalence   ratio  also   defines  the
saturation   temperature  (TDEW) .    If   we  were   to   use  the
 *  Note that equation  111-12  has VPHI in the  "J"  term as well.
In order to solve 111-12, we seeded a  computer  program with VPHI
and  iterated  111-12  until  the  seeded value and  the  calculated
value were essentially equal.

-------
   EO. VflPOR PHI VS COUP  PRESS
2.0

1.8

1.6

l.U

1.2

1.0

0.8

0.6
r
a.
o
Q.
cc
OD

_l
CC
0.2

0.0
  RMB«15.P U) oSSKPfl.
MVR00.65
N»65 _ _____

  RMSOO.P (1) eSSKPR. M
VR«0.65
  RM8«-IS,P (1) 08SKPR
.HVR00.65
N*6S

  RM80-35.P (1) 085KPR
.MVR00.6S
    0  500  1000  1500  2000  2500  3000  3SOO
       COMPRESSION  PRESSURE -KPfl (R8S)
                ,  Hi- I

-------
                             -15-

related values  for  TAD, TDROP,  and TDEW  corresponding to  a
given VPHI,  we  would find that  the  temperature  reserve would
be  essentially  zero.   But,  if  we   limit   the  saturation
temperature  (TDEW)  to  the  value  that  occurs  at  the  vapor
equivalence  ratio corresponding  to  minimum vapor  ratio (MVR)
required  to  initiate  combustion,   we  would  find  that  the
temperature  reserve  would begin to increase  once the value
TAD minus TDROP  exceeded  our .artificially  limited TDEW.  This
TRX value  is essentially  a  measure  of the equivalence ratio
reserve between  the  maximum vapor equivalence  ratio  that  can
be  sustained under the particular  pressure  and  temperature
conditions  after  evaporation,  and the  MVR  value.   If  we
truely wish  to  represent  the  temperature  reserve  (TR)  above
the MVR, then we  must also limit the temperature drop  (TDROP)
in equation 111-14 to that which would  occur  when evaporating
only  the  amount  of  fuel   necessary  to  achieve  the  MVR
selected.   Thus equation 111-14 becomes:

(111-15) TR* = TAD -  (TDROP @ MVR + TDEW @ MVR)

     Figure  III-3  is  graphic  depiction of  this  phenonon  for
several ambient  starting  temperatures   and  an  assumed minimum
vapor ratio  (MVR) of  0.65.

     Before  we   pursue   the   temperature   reserve   concept
further,  it  would be useful  to investigate the  effect  that
the assumed  minimum vapor  ratio  (MVR)  has on  the conclusions
that  may  be drawn  from  the model.   For  instance the lower
flamability  limit (LFL)  for  methanol at  atmospheric  pressure
corresponds  to   a  vapor   equivalence   ratio  of  0.455(19).
However,  Browning  (3)  suggests  that  an  MVR  of  0.63  is
necessary  for  initiation  of methanol  combustion.   Bardon (2)
on the other hand, suggests a ratio 0.61 is necessary.

     Figure  III-4 provides  an  insight  on  the effect  that  the
MVR has on the  temperature  reserve.   Trends are  plotted  for
MVRs  of   0.6   and   0.7,   and  for   two  ambient   starting
temperatures.   Note   that   the  leaner  the  assumed MVR,  the
greater the  temperature reserve.  Expressed  otherwise,  a high
*Note,  whenever  TR  was  less  than  or   equal  to  zero,  our
computer  program automatically  set  TR   equal  to  zero.   The
following figures  include this  assumption.   Future  work  may
find  the amount  of  negative  TR which  is  not  shown  in  our
figures  useful  in  determining the  amount of energy  required
to be  added to the system to  achieve  a  startable temperature
reserve.  However,  for  our  initial  understanding,  we  choose
to set negative TR values equal to zero.

-------
   EQ. TEMP RESERVE  VS  COMP  PRESS
400
                                             flM8*15.P moSSKPfl.
                                           HVR00.65
                                             RHBOO.P (1) «85KPfl,M
                                           VRC0.65
                                             flM8«-15.P (1) 085KPR
                                            .MVR00.6S
                                              flHB«-35.P (1) 885KPH
                                            .MVR00.65
                                            N-6S
        500  1000  1500  2000  2500  3000  3500
        COMPRESSION  PRESSURE -KPR (BBS)
              F/G.  W -

-------
   EO. TEMP RESERVE VS COMP PRESS
400
                                             RMBOIS.P (1) .085KPR
                                           .HVR80.60
                                             flM6015,P (1) .085KPR
                                           .MVRC0.70
                                             fi«8e-3S.P (1) ,«85KP
                                           fl.MVR«0.60
                                           N«6S

                                             flM8«-35.P (1) .085KP
                                           fl.MVReO.70
                                           N-6S
    0   500  1000  1500  2000 2500  3000  3500
        COMPRESSION  PRESSURE -KPfl (R8S)
              F/6.  UL -

-------
                             -18-
  ipor  equivalence   ratio   requirement   consumes  temperature
  'serve.   This of course  is due to the  leaner  mixture  having
   lower  saturation  temperature,  and to  less  sensible  heat
  i-inn r-£imi™nrar1  t- r\  Trar^/^iv-iTti  4-ViQ  dnallov  a m^Mi n 4-  o f  •Fnol    A 1 Q n
vapor
reserve.
a  lower  saturation  temperature,  and to  less  sensible  heat
being  removed  to vaporize the  smaller amount of  fuel.   Also
note   that   for  a  given   compression   pressure,   there  is
approximately   a   21°C  difference   in   temperature  reserve
between  an  MVR  of  0.6  and  an  MVR  of 0.7 at  either ambient
temperature.  However,  if we review figure III-3,  we observe
approximately   a  30°C   to   40°C  difference   in  temperature
reserve between different ambient  temperatures  (+15, 0,  -15*)
at  a  constant  MVR  (0.65).    Therefore,  the  selection  of MVR
can  have an  appreciable effect  on  the  results,   since  the
effect of the  value  selected can  be  a sizable  portion of the
effect due to a change in temperature with a constant MVR.

     The MVR selected for modeling cold  start performance may
be  a  function  of engine  combustion chamber  design  or it may
be  a   combination  of  factors  such  as   the  dryness  of  the
methanol,  the   chamber  design,  or  the  ignition  energy.   It
appears  that  one should  be  careful in  selecting  an  MVR for
modeling purposes, and  at least attempt  to be consistent when
possible  in comparing  data  from  other   investigations.   For
our analysis we choose  an MVR of  0.65 for  the  reason that at
least  two methanol  engines  with different  combustion chamber
designs  (Nissan NAPS-Z  and  Ricardo HRCC) have demonstrated
(19) (22)    that  they   can   be   operated   with    reasonable
performance at  this  overall  equivalence  ratio  in  a warmed-up
condition.  Also,  an  MVR of  0.65  is  slightly conservative
from the values used by Browning and Bardon.

     Now  that  the  concept  of   what  we  call  "temperature
reserve"   has   been  explained,  we  can  proceed  to  use  this
concept  to  evaluate  the  potential  limits  for  cold starting
with neat methanol.  Noticing  that the temperature  reserve in
Figure III-3  is  a  function  of  compression  pressure,  we can
compare  the temperature  reserve  at  the maximum  compression
pressure  with   one  engine  at a  given compression  ratio and
ambient temperature to  the  reserve for another  engine  with a
different   compression   ratio  or   at   a  different  ambient
temperature.   For  example,  in Figure  III-5 an  engine  with a
* The  difference  in temperature  reserve  between  -15°C  and
-35°C is approximately 40°C to 50°C.

-------
                             -19-

peak  compression  pressure* of  1000  kPa-abs** (130  psi  gage)
would  have  a  temperature  reserve  of   122°C  at  an  ambient
temperature  of   15°C.    That   same  engine   at   an  ambient
temperature  of  -35°C would  only  have  a  23°C  temperature
reserve.  If it were  assumed  that a 122°C temperature reserve
were necessary to  start  a methanol engine, from  Figure  III-5
the  maximum compression  pressure  would need  to  be  increased
to  around  2300 kPa-abs  (320  psi gage)***  to  obtain  a  vapor
equivalence ratio  of  0.65 under equilibrium conditions and an
ambient temperature of -35°C.

     Obviously increased  compression is  one means to increase
the  temperature  reserve  and  improve   the  fuel  evaporation
during   the  compression   stroke.    Another  means  is   by
increasing  the   throttling   during  cranking.   Because  the
volume  ratio  for  compression   would   remain  the  same  when
comparing wide open  throttle  (WOT)  to  closed throttle  (CT) ,
and  if  dynamic effects  were  ignored,  the  adiabatic pressure
ratio would remain the  same for either WOT or CT. Because the
pressure ratio  would  be  the  same,  the  adiabatic temperature
(TAD, Equation  111-12)  would remain the same for  either  WOT
or CT  (assuming  that  Tl  is essentially  the same  for both  WOT
and  CT  cranking  conditions  (ref.   18)).   However,  under  CT
conditions,   the  absolute  values of   p2  and  pi  would  be
lower  than  in  the  WOT case.   But,   because the  adiabatic
compression  temperature   (TAD)   remains  the  same,   a  re-
computation of  equation   111-12  and  111-14 with  the lower  p2
pressure would result  in a  higher  temperature  reserve than
would  occur under the  original WOT  conditions  because  the
partial pressure of the methanol vapor would be less.

     A  typical  4  cylinder  engine was  motored at speeds from
200  RPM to  600  RPM.   An analysis  of   the  vapor equivalence
ratio versus  ambient  temperature  (at  the  rounded  MAP values
from  Table  III-4)  was  shown  in figure  III-l,  and indicated
relatively  little  improvement in the ability to   pre-vaporized
the  fuel prior  to the   compression stroke.   However,  if  we
were  to  enter  these rounded-off  results  into  equation 111-13
for  the  "MAP"  term,  and  recompute  the  temperature reserve by
*    Assumes  that  measured  cranking  compression  adequately
     represents the actual "pz" in the P-V cycle.
**   This  pressure is  assumed  to  be  rather  typical  of  low
     compression SI engines converted to run on methanol.
     Our current  test engine, a  2  liter  Nissan NAPS-Z engine
     with 12.8:1 compression  ratio  has  a  cranking compression
     of  around  1960 kPa-abs  (270 psi gage)  when motored on a
     dynamometer  at  200  RPM,  note  the  normal  in-vehicle
     cranking speed  for  this engine is specified as  300 RPM.
** *

-------
   EQ. TEHP RESERVE VS COHP PRESS
400
                                             RHB«15,P (D085KPR.
                                           MVR00.65
                                           N-65    _ ___

                                             RMBOO.P (1) eSSKPR.M
                                           VR00.65
                                             RHBO-15.P (U Q85KPR
                                           .MVR00.65
                                           N«65

                                             RMBO-35.P (1) «85KPR
                                           .HVReO.65
                                           N-65
       500  1000  1500  2000 2500  3000 3500
       COMPRESSION  PRESSURE  -KPR (BBS)
              F.,
in- -

-------
                             -21-
the  method  in  equation  111-15,   we  would  find  that  the
temperature  reserve  would increase  more  markedly at  the  low
temperatures.  Figure III-6,  suggests that  for  an engine with
1000  kPa  maximium  cranking   pressure and  85  kPa  MAP  ,  the
temperature  reserve   at   15°C  ambient  temperature  would  be
increased by about 35  percent  by lowering the MAP from 85 kPa
to  65  kPa,  but  at  -35°C  ambient,  the  reserve  would  be
increased by almost   160  percent over  the  85  kPa  value.   In
fact the temperature  reserve  of 60°C  at -35°C  ambient  and 65
kPa is  almost  50 percent  of  the reserve at  15°C ambient and
85  kPa,  a  condition  at  which  most neat  methanol  engines
should start.

                          Table. III-4
     RPM

     200
     400
     600
 Cranking MAP'

MAP*

91.72
75.53
62.86
Round-off**

    90
    75
    65
     curve fit  (R2  =  .995),  MAP =  105.5567  -  (. 072125) (RPM)
     300 RPM = 84 kPa rounded to 85 kPA.
     Round-off represents  arbitrarily  picked  numbers close to
     the observed values for analysis purposes.
     Barometric Pressure = 99.17 kPa (wet).
     It should  be noted, however,  that  this analysis assumes
that  some means  is  used to  maintain  the maximum  cranking
pressure  at  a constant level when  lowering  the  MAP.   If such
means  were   not   employed,   then   the   maximum  compression
pressure would  be  lowered by the ratio of  the new  MAP to the
old  MAP,  unless  other  factors   such  as  leakage  during  the
compression   stroke   would  limit   the   maximum  compression
pressure at the higher MAP condition to a  value  that  is below
that which  could be achieved at  the lower MAP  condition with
no  leakage.   In comparing the measured  p2  in  our engine  at
the  different  cranking speeds to that at  200 RPM,  the actual
p2  is  about  7.5  percent higher  at  400  RPM   and   about  12
percent higher  at  600 RPM.   For  a large  V-8 gasoline truck
engine these  values  are  about 10 percent  at 400 RPM  and  18
percent  at   600  RPM.   Because  these values  are  relatively
small,   in this  analysis  we will   simply  assume  a  straight
ratio of  the new MAP  value  divided by the  MAP  at  the normal
cranking  speed  to  be  multiplied  by  the  pz  at  the  normal
cranking  speed.   Those  values  will then  be  rounded  to  the
nearest 50  kPa  which  will  be  designated  as  "MAP  corrected
pz" or "corrected p2".

-------
     EQ. TEMP RESERVE  VS  COUP  PRESS
  400
  350
u
UJ
o
  300
  250
oc
UJ
2200
ec
UJ

3
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                                                RH3015.P (1) 085KPH.
                                              MVR00.65
                                                RM801S.P (1) 06SKPP,
                                              HVR00.65
                                                ftM8o-3S,P (1) 085KPP
                                              .MVR00.6S
                                              N>65
                                                RM8e-35,P (1)
                                              .MVR00.6S
                                              N-69
         500  1000  1SOO  2000  2500  3000  3500
         COMPRESSION  PRESSURE -KPR (BBS)
                    . nr -

-------
                             -23-

     Reconsidering  the  lower MAP,  if  there  were a  means  to
substantially  reduce the  cranking  MAP  to  the  vicinity  of
around 30 kPa  -  absolute  (20.5  inches Hg manifold vacuum),  we
could essentially  raise the  temperature reserve  from  23°C  to
179°C  at  -35°C   ambient  if  we  could  maintain  a   1000  kPa
maximum  cranking  pressure  (i.e.,  57°C  higher  than the  value
for 85 kPa MAP at 15°C  ambient  - see figure III-7).   However,
if  we  could  not  maintain  the  maximum  cranking   pressure
constant,  and the  maximum  p2  was  lowered to  350   kPa  (ie.
MAP  corrected PZ),  the  temperature  reserve  would  be  only
46°C  —  76°C  lower than the reserve  at  the  15°C ambient base
condition.  Achieving  a 50  kPa  cranking  MAP  might  be  more
practical.*    A   50  kPa   cranking  MAP  with  constant  peak
pressure  would  still  allow  a   97°C  temperature  reserve  at
-35°C  ambient, nearly  equivalent to  the  temperature reserve
of  122°C  at  a  normal  cranking  speed  and  15°C   ambient.
Reducing  the  peak pressure  by the MAP ratio to 600 kPa, would
lower the reserve to just 38°C.

     Compared  to the 23°C  reserve  for  85kPa  MAP and  -35°C
ambient  (see   figure III-7),  the  38°C  reserve at 50  kPa  and
the  46°C reserve at 30kPa  do   show  an  improvement  in  the
temperature   reserve  with   lower  cranking   MAP  values.
However,   lowering the  ambient  temperature causes a  decrease
in  the   temperature  reserve,    and   this  decrease   is  not
compensated  by  the MAP  effect.   Therefore,  a  substantial
amount of the normal temperature reserve at  15°C (e.g.,  the
difference between 122°C and 46°C) would be  presumably needed
to  be made-up by other means  if the engine  were  to start.
Further,   achieving  very   low   cranking  MAPs  with  practical
cranking  hardware  while maintaining  sufficient control of the
fuel-air mixture  could  be  difficult.   Also,  smooth transition
from  a high vacuum crank  condition to a normal idle  condition
might be  tricky.   Even  so,  the  effect of lower  manifold  and
cylinder  pressures  should  not be over looked.   For a consumer
acceptable  vehicle,  the  effect  of  altitude  versus  lower
cylinder  pressure  may  affect  the  starting  capability,  and
should also  be  given  consideration  (i.e.,  the  5 inches  Hg
pressure  differential  due  to   an  altitude  change  of  1524
meters -5000 ft.- is approximately 17 kPa).
* 50  kPa  can be achieved by increasing  the  cranking  speed in
our  engine  to  approximately  770  RPM  (normal  speed is  300
RPM),  or  presumably  by  increasing the  throttling  at  the
normal cranking speed.

-------
   EO. TEHP RESERVE  VS  COMP  PRESS
400
                                              RMBOIS.P moSSKPfl.
                                            HVRoO.65
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                                              flMBo-3S.P (1)030KPR
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                                            N-58
    0  500  1000  1500  2000  2500  3000  3500
       COMPRESSION  PRESSURE -KPflfflBS)
               F/6.  zzr-7

-------
                             -25-

     In  discussing   means   of  increasing   the   temperature
reserve,  we   have   digressed  somewhat   from   the   original
hypothesis of comparing the temperature  reserve  (TR)  at  known
starting conditions  to those  at unknown cold conditions.   It
is  generally  accepted  that  the lower  temperature limit  for
unassisted  starting  of gasoline engines converted to run  on
methanol  is  around  10°C  (50°F)  (10)(11).   Without  changing
pistons  such  engines  usually have  rather  low  compression
ratios  (in the  range of 8:1  to 9:1).  A cranking  compression
of  lOOOkPa-abs  (130  psi  gage) was  earlier  assumed  as  being
reasonably  representative  of  such  engines.    Under  these
conditions  (10°C  ambient  and  1000  kPa  compression pressure),
the temperature reserve would be around 112°C  for an MVR  of
0.65 and a MAP of 85 kPa (see figure III-8).

     Gardiner et.  al,  (10)  extended  these  limits  somewhat  by
achieving successful cold starts at 0°C.  The engine  used  was
a 2 cylinder  air  cooled 4 cycle SI engine  with  a compression
ratio  of  5.84:1.   The  manufacturer  suggests   that  normal
compression  pressure  on  this  engine is  around  750  kPa-abs
(94psi  gage).   For  this pressure  and assuming  a  65   kPa  MAP
(since   Gardiner's   cranking   speed  was   600    RPM*),   the
temperature  reserve  at 0°C  ambient  with a  non-corrected  p2
is 96°C, a  value  not substantially different from the assumed
standard engine at  112°C  temperature  reserve.    The  reserve
with   a   corrected   pz  of   550   kPa   is   60°C.   A  visual
comparison  of  Gardiner's engine  to  the  standard engine  is
shown in figure III-8 as well.

     M.A.N., on the  other hand, with a  modified  diesel  (CI)
engine with spark assist has  reported instantaneous  starts  at
-2Q5C— (^)°F-ref. 12) .   Informal  accounts  (15)  indicate  3 second
starts  at  -32°C.    The exact  cranking  compression   of  the
M.A.N.   engine    is   unknown   to   us,    but  from   informal
conversations with M.A.N., a  range  of 25  to 30  bar  (2500  to
3000  kPa)  was  suggested.    Typically,   CI   engines  are  not
throttled,  and therefore,  the cranking MAP  would  be closer  to
atmospheric  pressure,  or  approximately  100  kPa  for   this
analysis.  Assuming that the M.A.N.  engine has a  100  kPa MAP,
the temperature reserve at  -35°C ambient would  range between
107°C to 131°C  (see  Figure  III-8)  for the  range  of  estimated
compression pressures,  values  which  are  still at or above  the
value need to start the standard SI engine at +10°C ambient.
* The  normal  cranking speed  for  this engine  is between  200
and 400  RPM.   A nominal  value  of 300 RPM and 85 kPA  MAP was
assumed for this analysis.

-------
                             -26-

     If some throttling were used, for instance to  achieve  85
kPA MAP  during crank,  the  corrected MAP  temperature  reserve
(approx.  2100  kPa-p2)at -35°C ambient for the low  range  (25
bar) would  increase  from 107°C  to  110°C.   The reserve  would
increase approximately 2°C for the higher  value  (approx.  2500
kPa corrected  p2) as  well  -  see figure  III-9.   If we  were
to decrease the MAP even further to 65 kPa, we would increase
the temperature reserve  another  7°C  at  the low range and 14°C
at  the high range (117°C  and  147°C  respectively).   In  fact,
the temperature reserve  for  -50°  ambient  of 109°C at the high
compression pressure and  65  kPa  MAP  is  nearly sufficient  for
starting at -50°C  since the  reserve  of  our standard SI  engine
at 10°C ambient is only 112°C.

     These  results  of  increasing  temperature  reserve  with
decreasing   MAP   and   corrected  p2   match   our   earlier
observations with  one  exception.  The exception  is, in  this
case decreasing  the MAP causes  a  more  observable  change  in
the  projected  starting  capabilities  at  the  lower  ambient
temperatures.    The  key to  the  more  observable  change  lies
simply in the fact that  the  high  compression  pressures  result
in  temperatures   reserves  that  are  initially  nearer to  the
reserve for our standard engine at 10°C.

     Table  III-5  provides  a  handy comparison  of  our analysis
for the  three engines  that  have  been discussed,  the mythical
standard  SI engine,  Gardiner's  engine,   and  the  M.A.N.  CI
spark assisted engine  to projections  based on the Temperature
Reserve concept.  High  compression pressures  and/or  a  low MAP
during  cranking   appear to  be avenues  to achieve  potential
unassisted  cold   start  capabilities   in  the   -30°C  to  -40°C
range  --  the  high  compression  through  higher  in-cylinder
temperatures  during  the  cranking  cycle,   and  the  high  speed
cranking  through   reduction  of  the  partial  pressure of  the
fuel due to the lower MAP.

     The  high  speed  cranking  may  have   an additonal  factor
that assists  in  the cold starting phenomonon.  With the high
cranking   speed,   there  exists   the  potential    for   more
compression strokes  to occur  during cranking  prior  to  start.
Since  the  actual engine cranking  is  not  adiabatic,   each
cranking stroke  leaves behind some  residual  fuel vapor  at  a
temperature .higher  than  the  ambient   mixture,   potentially
requiring  less heat  (or  lower   temperature  reserve) on  the
next stroke to  achieve  the  MVR  (minimum vapor  equivalence
ratio)  necessary  for  starting.   Gardiner's starting procedure
of  cranking the  engine  at  600 RPM for 10 seconds  (i.e.,  50
compression strokes) and  then  resting  for  10 seconds followed
by  two  replications   of   this   cycle,   allows   up  to   150
compression strokes  to occur  for what Gardiner accepted  as  a

-------
               -27-
     EO. TENP RESERVE VS  COUP  PRESS
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-------
                             -28-

                            Table III-5



                       Actual vs. Projected

                       Starting  Capabilities
          Condition
Reference
   -Standard SI
   -Gardiner (600 RPM crank)
   -M.A.N. (12)

   -M.A.N. (15)
Projected***

   -SI engine with high
    speed cranking (770 RPM)
   -High compression SI engine
    0 w/norrnal cranking (300 RPM)
                              Max Cyl
                             Press(kPa)*
      TR   Ambient
       >C)   Start
1000
550
2500
3000
2500
3000
85
65
100
100
100
100
112
60
142
173
107
131
10°C
0°C
-20°C
-20°C
-32°C
-32°C
                                  600
    0 w/high
-39°C
   -M.A.N
        speed cranking  (770 RPM)
   -M.A.N.
    during
      projected capabilities

      w/intake throttling
      crank
50
112
- 2°C
2100**

2500
3000
1600
2000
85**
1200
100
100
65
65
112
50
112
112
112
112
-34°C
112
-33°C
-43°C
-37°C
-49°C
* Corrected Pz
**Max  pressure  and  MAP  data  from current  test engine,
NAPS-Z with 12.8:1 compression  ratio.
***Ambient  Start temperatures  are  linearly  interprolated
computer output matrix based on:
AMBN
ter output matrix based on:

= AMBL -  ([AMBL-AMBH][(TRS-TRL)/(TRH-TRL)])
.ere:
f R.. = n *=*T*T  amHiont- c i- a K f- -f-omnoKaf-iiyd
                                                       a  Nissan
                                                       from  the
   Where:
   AMBN = new ambient start temperature
   AMBH = ambient temperature with a TR above TRS
   AMBL = ambient temperature with a TR below TRS
   TRS  = the TR for the standard SI engine at 10°C ambient,  85
            kPA-MAP, and 1000 kPa = 111.955°C
   TRH  = TR corresponding to AMBH
   TRL  = TR corresponding to AMBL

-------
                             -29-

successful  start.    A normal  cranking  speed  (200-300  RPM)
would have  only 30 to 50 percent as  many  compression  strokes
for  a  given  time.   Thus if  the exhaust  residuals which  in
this case  are  fuel  vapors, build  up with  successive  engine
cycles,  the  high speed  cranking  could have  an advantage  in
this  area  as  well,   and  may  be  partially responsible  for
Gardiner's ability to  start with a  lower  temperature  reserve
than  the  value  predicted  for   other  investigators  with  a
so-called standard engine (see figure III-8 and Table III-5).

     Detroit   Diesel   Allison   (DDAD)   utilized   a   similar
approach  with  exhaust  residuals   to   develop   successful
autoignition   in  its   methanol   bus   engine   (13).     The
autoignition  temperature  of methanol was  identified by DDAD,
and  the temperature  loss  due  to  methanol  evaporation  was
calculated.     DDAD   then  increased  boost   pressure   and
temperature, and varied exhaust back  pressure  to  increase the
exhaust  residuals  in the  engine  in  order  to  recover  the
original  temperature.   Quite  possibly  a  butterfly valve  in
the  exhaust  pipe  similar  to  those  used  for  early  fuel
evaporation  systems  (EFE)   on  V-8  gasoline  engines,  when
applied  to   4  cylinder  methanol engines  would,  when  closed
during cranking, increase the  residual  fuel  vapor  retained in
the  cylinder   from   successive  cranking   cycles.    Another
possibility would  be  to  delay  the fuel (which is possible on
electronically  fuel   injected  engines)  until  after  several
cranking  revolutions  in   order  to  have   hotter  cranking
residuals.

     Since   Table   III-5  suggests    the   potential  starting
capability  in the -30  to  -40°C range  for  high  compression
ratio engines,  and since there is not  a  plethora  of methanol
engines with  -30  to  -40°C starting  capability,  other  factors
need to  be  examined   (Note  also that most  of  the  cold start
studies reported  in  the  literature  have  generally  been  with
relatively   low  mechanical   compression   or   low  cranking
compression engines).  First the  analysis  presented is  a  bulk
equilibrium  analysis,  and is  intended only  to  show the outer
limits of potential  start capabilities (i.e.,  if  equilibrium
conditions   cannot   support  a   combustible   vapor  mixture,
non-equilibrium conditions  do  not stand much of a  chance for
starting).   In  reality,  the evaporation  process  which  occurs
during  the  compression cycle  is  a  series of  non-equilibrium
processes that  occur during the  finite time of  the piston's
upward  travel.   Time is  a  factor,  in  heat  transfer to  both
the  methanol  droplets  and  to  the  surrounding  structure.
Compression  heat  is  continuously transferred  to  both  at  a
rate depending  on the difference in  temperature  and the  heat
transfer  coefficient   for  each.    The  maximum   allowable
evaporation  rate  of  methanol   plays   a  critical  role  in
determining if  a  sufficient amount  of fuel can  be  evaporated
by the end of the compression stroke.

-------
                             -30-

   It  is   apparent   that  increasing   the   compression  pressure
increases  the  temperature  reserve  which   should   increase  the
potential  for successful  starting  at  cold  ambient  temperatures.
Slightly   less   apparent   is   the   observation   that  when   the
temperature   reserve   is   the   greatest  (i.e.,    normal   summer
temperatures, 25° - 40°C),  the  portion  of the initial  liquid fuel
that can  be vaporized in the inlet  manifold  (based  on equilibrium
conditions)  is  also  the  most  favorable  (see figure  III-l)  thus
tending to  prevent  the evaporation  rate  of methanol  from  being a
potential limiting factor because  the amount that  is  prevaporized
can  be  a  sizable  portion  of   the  MVR  (20%-30%) .   Both  factors
(increased   temperature   reserve   and  increased   manifold   vapor)
enhance  starting  capabilities.   Conversely,  when the  temperature
reserve is the smallest (i.e, low  ambient  temperature),  the amount
that can  be  vaporized in  the inlet manifold is also minimal, which
indicates  low  ambient temperature   conditions  have  two  factors
working  against  successful  cold starts  -  (1) reduced  ability to
vaporize  the fuel  during the compression stroke, and  (2)  reduced
capability  to  pre-vaporize  some fuel in  the  intake  manifold prior
to the compression stroke.

   To  further explain  the  failure   of  real  methanol   engines  to
start  at  low ambient  temperatures  for  which   their  theoretical
temperature  reserve  seems  adequate  based on  the  testing   of  low
compression  engines  at   higher  ambient  temperatures  (10+°C),  we
must   examine  how   ambient   temperature   itself   effects   the
non-equilibrium processes which  are  at  work.,- One fact which seems
important  is  that  as  the  ambient  temperature  decreases,  the
difference   between   the  peak   compression   temperature  and  the
ambient  temperature  also decreases.   This  temperature  difference
is  roughly  the driving force for  initial  heating of fuel droplets
which began  at ambient temperature.   Later,  when  fuel  droplets can
be  approximated  as  having  been  warmed  to  the  boiling  point
corresponding  to  the  total pressure  in the cylinder,  the  driving
temperature  difference  for  heat transfer  will  also be  reduced at
lower ambient  temperatures  since the peak  compression temperature
is  reduced  but  total  cylinder  pressure  is  not.   Thus,  heat
transfer  and evaporation  will  be  slower  for  a   given  temperature
reserve at lower ambient temperatures, impairing  starting ability.

   In  fact,  the  evaporation  rate  of  methanol  may  be the  only
extremely critical factor in starting a methanol  engine  under cold
ambient     conditions.      When   .  evaluating      the     M.A.N.

-------
                             -31-

system*  (12),  it is  interesting  to  note  that the  fuel  is
injected  late  in  the compression  stroke.   This  provides  a
very large temperature difference between  the  air  temperature
and the  fuel  temperature which enhances heat  transfer  to  the
fuel.    Secondly,  and  possibly more  important  is  that  the
engine design  does  not attempt to  evaporate all  of  the fuel
at once.   Instead,  the fuel  is deposited  in  a  bowl  in  the
piston.   There,  only  the  exposed  surface   is   open  for
evaporation,  and  the  evaporation  that does  occur, cools  the
piston,  not  just the  air.   After  ignition  by a  spark plug,
the flame provides  for controlled evaporation  and combustion
of the  remaining fuel in a  manner  that allows an acceptable
pressure rise in the cylinder.

     Gardiner,  on the  other hand,  indicated  that  increasing
the  amount  of  liquid  fuel to the  engine  from  an  initial
liquid equivalence  ratio  of about  1.0  to  about 3.0  improved
the  cold  starting  performance of  a carbureted  SI  engine.
However, increasing the  initial liquid  ratio from 3.0  out  to
10.0  made  little  change in the  starting  performance.   One
explanation of  this may be  that  the additional fuel surface
area  achieved  by an  equivalence  ratio  of 3.0 was  enough  to
prevent  the  evaporation rate  from being  a  limiting  factor,
such  that  the  work  of compression was the  limiting  factor  as
equivalence ratio was  further   increased.   A given  amount  of
compression  work  will  support  only  a  certain  amount  of
methanol vapor at equilibrium  conditions for  a  given ambient
temperature  (see  figure  III-2).  Another  explanation is that
the evaporation  rate was limiting throughout Gardiner's range
of equivalence ratios, but  because  of poor fuel atomimization
and distribution, the  fuel   surface  area in the cylinder  did
not  increase  past  an  equivalence  ratio of  3.0.   Gardiner's
own explanation  is  also plausible,  namely that surface area
and  evaporation   rate  did  steadily  increase  with increasing
equivalence  ratio,   but   (expressed  in   our  terms),   the
temperature reserve was  possibly  consumed  by the need to heat
large amounts  of  liquid  fuel to the  boiling  point.   Gardiner
does  not indicate  whether  he  attempted operation  beyond  an
initial  liquid equivalence ratio of  10.0.

     From this information,   a  hypothesis is put forward that
may be  difficult to justify, but none the  less may stimulate
further  thought  in  this  area.   It is hypothesized  that under
cold  starting   conditions   that the  rate  of   evaporation  of
methanol  is  the limiting   factor.   During  the  compression
stroke,  it  is  hypothesized  that  the gas  temperature remains
*A modified diesel engine with spark assist.

-------
                             -32-

well  above   the   saturation  temperature   for   the   partial
pressure of  methanol;  but because  rate  limitations have  not
allowed sufficient methanol  to evaporate,  the  temperature has
not been driven down.   Figure 5  in Bardon's paper (2)  which
shows   the   measured   cylinder  temperature   after  methanol
vaporization  to   be   much   greater   than  the   theoretical
temperature   after  vaporization   tends   to   support   this
hypothesis.   If  sufficient  methanol  has  not  evaporated  to
drive the  temperature  down,  also sufficient methanol  vapor is
not available to form a combustible mixture.

     Increasing  the  overall  liquid  mixture   in  order  to
potentially  increase  the surface  area of the  fuel available
for  heat  transfer  and subsequent  evaporation  (such  that  a
combustible  mixture   is   formed)  would   require  that  the
remainder of this  excess  liquid  fuel must be consumed  by the
flame.  However,  if the  excess   fuel  to  be  consumed by the
flame was  delivered to  the  combustion  chamber  in a  manner
such   that   large  droplets   remain  after   the   initial
vaporization,  there  is  evidence  to  suggest  that  the  flame
will   not   propagate    due   to   insufficient   preliminary
volatilization  of  the  large  droplets.    Hence,  the  initial
flame  front  could  essentially  be  quenched  by  the  large
droplets  resulting from  the excess  fuel  causing  a  no-start
condition.

     Burgoyne  and  Cohen  (20)  conducted  an experiment  with
tetralin   and   air  mixtures  which   evaluated   the   flame
propagation  of  various  sized  fuel   droplets  surrounded  by
just-flammable  fuel  vapor.   Their  apparatus  was  configured
such that  the  droplets flowed downward through a burning zone
at  the  end  of  a  long  tube.   The combustion was  supported by
the  just-flammable  fuel  vapor,  but the  vapor  flame would not
propagate  up  the  tube.   When  the  droplets  were added  in
appropriate  amounts  the  flame then  propagated  up the  tube.
They  found  that  for droplets  above 300 microns  in diameter,
the  flame  would  not propagate upwards  against  the  flow which
they suggest is due to  the velocity of  the fall  being greater
than  the  burning  velocity  such  that  the mass motion  of the
droplets  prevents  propagation.   Such  mass  motion  is  always
evident  to  some  degree  within  the  cylinder  during  the
compression  stroke.   The  spark  plug  could be  envisioned as
the  location  of  the   just-flamable  burning  zone  and  the
droplets would pass  through this  region  due to the  swirl in
the  cylinder.   This  might  suggest that  the more quiescent
combustion   chambers   need   less   temperature   reserve  for
starting than one  with high swirl  even though the high swirl
may   be  more   desirable   once  the   engine  is   running.
Conversely,  the  higher swirl engine may need  a  larger  MVR to
compensate the initial  vapor  flame  speed  for  the higher local
air velocities.

-------
                             -33-

     Another  aspect  of  Burgoyne's  and  Cohen's  work is  the
concept  of  two  regimes  of  droplet  evaporation  prior  to
combustion.   The  two  methods are  evaporation controlled  by
diffusion, and evaporation  controlled by heat  transfer.   The
boundary  between  the  two  mechanisms is  controlled by  what
they define as the  "stabilized  temperature"  which  is  defined
as   the   "   'wet-bulb'   temperature  which   is    below   the
boiling-point  by  reason  of  the  heat  carried  away  by  the
vapour".    If   we  compare   the  heat  of  vaporization  between
methanol  (474  BTU/lb  or  1.102  J/kg -  ref.  4)   to  that  of
tetralin  (0.201  J/kg  -  ref.  23),  we see  that the  amount  of
heat that  can be  absorbed  by   the  heat  of  vaporization  is
about  5.5  times  greater  for   methanol  than  for  tetralin.
Therefore, one would  expect the "stabilized  temperature"  for
methanol to  be much higher than that for  tetralin because  of
the  additonal  heat  transferred  away  from  the droplet by the
evaporating methanol.

     Burgoyne  and   Cohen   also  point  out   that  aerosols
(droplets  under   10  microns)   have  much   higher   burning
velocities  because   the   time  required   to   evaporate  the
droplets  is  much shorter  than  for  larger droplets.   If  we
assume that  droplet  evaporation will occur much  more quickly
in the heat transfer  controlled  regime  than in the diffusion
controlled  regime,   then  because   of  the  higher  stabilized
temperature  for  methanol  which keeps   the  methanol  in  the
diffusion  controlled  regime  longer,  one  would  expect  the
pre-flame  evaporation  of  methanol  droplets  to  take   much
longer than tetralin.

     Transferring   these   methanol   pre-flame   evaporation
assumptions back  to  Burgoyne's  and  Cohen's propagation  tube
results,  we  would  predict  that  the 300 micron upper  droplet
size  limit for  propagation  with   tetralin  in  air  mixtures
would be  reduced  (possibly  substantially)  for  methanol in air
mixtures  because  of  the  longer  time  needed  for  pre-flame
volatilization of  the methanol  droplets.   The lack  of  flame
propagation due to the pre-flame evaporation  rate  and droplet
size factors could explain  the  failure  of  other experimenters
to achieve  cold  start temperatures  as  low as  predicted  here
through  purely liquid enrichment.   To  be  accurate,  however,
it must  be pointed  out  that  not   enough  experimenters  have
reported  cold  start  investigations  with  high   compression
ratios  or  low  MAPs  to   allow   a   fair   appraisal   of  the
limitations of unassisted methanol  cold  starts.   Furthermore,
some experimenters have added  liquid fuel  in a way which may
not have increased the surface  area available for evaporation
(the additional  fuel may have merely pooled  in  the  manifold
or have merely formed  a  thicker  layer on the cylinder walls),
or  added  fuel  in  a  way  that  the  fuel  droplets  may  have
congealed  forming  droplets  too  large  to  allow  pre-flame
volatilization and subsequent flame propagation.

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                             -34-

     Based  on  this  hypothesis,  CI  engines  and  modified  CI
engines  (M.A.N.)   are  expected  to  be  easier to  start  than
homogeneous charged SI engines.  There are  several  factors  to
support this statement.  First,  CI  engines  almost  always have
higher  compression ratios  which  have correspondingly  higher
compression  temperatures.   Also, CI  engines inject  the fuel
late in the  compression  stroke where  there is  a much  larger
temperature  difference  between  the  air  and  the fuel  (which
should increase the vaporization  rate).  Further,  because the
fuel  is  injected  by  a  high  pressure  injection  system,  it
usually has smaller droplets than a  homogeneous  charge  engine
which  enhances  both  diffusion  controlled  evaporation  and
pre-flame  heat  transfer  controlled evaporation.   Lastly,  the
initial part  of the  injection  cycle  creates a  very  lean f/a
ratio  which  also  improves  partial  pressure   or  diffusion
controlled evaporation.  All four  factors,  higher  compression
temperature,    larger   temperature    differential,    smaller
droplets,   and  lean  f/a  ratio   enhance  the  evaporation  rate.
Because the  potential  for  the  evaporation  rate  during  the
initial part of the  injection  cycle is much greater under the
CI type conditions  than  those   in  an  SI  engine,  less overall
liquid  fuel  should  be required  to  achieve  a  combustible
vapor.  Therefore, since less overall  liquid  may be required,
the probability of excess  liquid or  large  droplets quenching
the  initial  flame   front  would   most  likely  be  reduced.
Finally,   even  if   excess   liquid   does  occur,  certain  CI
combustion   systems   (e.g.,   M.A.N.)  allow  for   controlled
evaporation  of   the  excess   liquid  in   order   to   prevent
quenching of the pilot flame.

     Since many of the features  that make  cold starting more
probable  with  a  CI  or  late  (near TDC)  injecting  style  of
engine  may not be easily transferred  to   an SI engine,  the
outlook for  the cold  start  performance of  homogeneous charged
SI  engine  does  not  look  promising.   Items   that  may  be
critical  to  cold  start performance of the  SI methanol  engine
without starting  aids  include  devising  means to increase the
evaporation   rates   through    such   mechanisms   as   higher
compression   temperatures    through   increased    compression
ratios, smaller droplet sizes  through  improved fuel systems
(e.g.   EFI,  sequential  EFI,  higher pressure EFI,  or  other
atomization  systems),  higher   cranking  speeds  or  increased
throttling, reduced heat transfer to the structure  during the
compression  stroke,  increased   heat  transfer between the air
and  the  fuel  during  the  compression   stroke,   and  possibly
devising  means  to better use  the cranking  residuals.   Flash
vaporization might be another possibility as  described  by Oza
(14),    however    this   technique  will   be  discussed   more
thoroughly in the  chapter on physical vaporization.

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                             -35-
     C.  Summary

     In  summary,  we  have  identified  several  instances  of
successful  cold  starting  of  unassisted   methanol   engines
(Gardiner  (10)  at   0°C,   and  M.A.N.   officially  at  -20°C,
unofficially at -32°C).  A  hypothesis  has  been developed that
suggests that the limiting  factor in unassisted  methanol cold
starting is not the equilibrium vapor  pressure curve, but the
rate and time sequence of the  fuel  evaporation.   Experimental
work  by  Gardiner  (10),   and  modeling  by  Bardon  (2)  and
Browning (3) tend to support this  hypothesis.

     This  information  then suggests  that   the potential  for
cold starting   a  methanol  engine without  assist at  very low
ambient   temperatures   is   still    a   possibility.    High
compressions  ratios  and  smaller  droplets  are  two  possible
factors   required   to   achieve    acceptable   cold   start
capabilities.    A   list   of    perceived    advantages   and
disadvantages is given below.
     1.   Advantages

         a)  Cost  of  Vehicle Fuel  System -
             cost  of  a
             unassisted
             expensive  than
             associated  power
             single-fuel system
             systems.
                     Most  likely
 single-fuel  vehicle fuel  system
methanol  cold  starting will  be
     one  with  auxiliary  heaters
        requirements.   Certainly
        should  be less than dual
                 the
                 for
                less
                 and
                 the
                fuel
         b)  Cost  of  Fuel  -
             additives  should
             volatility additives
       Methanol
       cost  less
with  no  volatility
 than methanol  with
         c)  Emissions  -  With  a  single  fuel,   it  will  most
             likely  be  easier  to  measure,  characterize,  and
             control  both  exhaust  and evaporative  emissions
             compared  to  methanol  with volatility  additives
             or dual fuel systems.

         d)  Enforcement - With  a  single fuel,  disputes  over
             using or selling  a  winter  fuel during the summer
             or in an unacceptable region of  the country will
             not occur.

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                        -36-

    e)   Vehicle Maintenance - A single-fuel vehicle  fuel
        system does not need  the range of  authority  that
        a system would need  for methanol  with volatility
        additives.    This  could  make  the   single-fuel
        system less complex  and easier to  maintain.   The
        single-  fuel  system  would  almost  certainly  be
        less complex than dual fuel  systems.

2.   Disadvantages

    a)   Universal Demonstration - With  the exception  of
        a  few   situations,   universal   application   of
        successful   cold  start  technology has  not  been
        demonstrated.   No  successful  cold start with  SI
        engines   in   the   -35°C    range    have    been
        demonstrated.

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