Dimethyl ether autoignition in a rapid compression machine_ Experiments and chemical kinetic modeling.pdf

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1. Dimethyl ether autoignition in a rapid compression machine:
Experiments and chemical kinetic modeling
Gaurav Mittal
a
, Marcos Chaos
b,
⁎, Chih-Jen Sung
a
, Frederick L. Dryer
b
a
Department of Mechanical and Aerospace Engineering, Case Western Reserve University, Cleveland, OH 44106, USA
b
Department of Mechanical and Aerospace Engineering, Princeton University, Princeton, NJ 08544, USA
ARTICLE INFO ABSTRACT
Article history:
Received 15 February 2008
Received in revised form 27 May 2008
Accepted 29 May 2008
Dimethyl ether (DME) autoignition at elevated pressures and relatively low temperatures is
experimentally investigated using a rapid compression machine (RCM). DME/O
2/N
2
homogeneous mixtures are studied over an equivalence ratio range of 0.43 –1.5 and at
compressed pressures ranging from 10 to 20 bar and compressed temperatures from 615 to
735 K. At these conditions RCM results show the well-known two-stage ignition
characteristics of DME and the negative temperature coefficient (NTC) region is noted to
become more prominent at lower pressures and for oxygen lean mixtures. Furthermore, the
first-stage ignition delay is found to be insensitive to changes in pressure and equivalence
ratio. To help interpret the experimental results, chemical kinetic simulations of the ignition
process are carried out using available detailed kinetic models and, in general, good
agreement is obtained when using the model of Zhao et al. [Int. J. Chem. Kinet. 40, 2008, 1–
18]. Sensitivity analyses are carried out to help identify important reactions. Lastly, while it
is implicitly assumed in many rapid compression studies that chemical changes from the
initial charge conditions that might occur during compression are negligible, it is herein
shown with the help of Computational Singular Perturbation (CSP) analyses that chemical
species formed during compression with little evolved exothermicity can considerably
affect autoignition observations. Therefore, it is essential to simulate both compression and
post-compression processes occurring in the RCM experiment, in order to properly interpret
RCM ignition delay results.
© 2008 Elsevier B.V. All rights reserved.
Keywords:
Dimethyl ether
Ignition
Rapid compression
Chemical kinetics
Induction chemistry
1. Introduction
Environmental pollution, energy security, and future oil
supplies are concerns that have driven the global community
to seek nonpetroleum-based alternative fuels, along with
more advanced energy technologies to increase the efficiency
of energy use. Dimethyl ether (DME) appears to have a large
potential as an energy source. DME has been proposed as a
promising alternative to diesel fuels due to its high cetane
number[1]. In recent years DME has emerged as a fuel additive
to reduce particulate and NO
xemissions due to its overall low
sooting and polluting potentials. DME can be mass-produced
from extensive feedstocks, including natural gas, coal, and
biomass; its physical properties are similar to those of
Liquified Petroleum Gases (LPG), so DME can be stored and
distributed using established LPG infrastructures[2,3].In
addition, DME can also be used as an ignition enhancer in
propulsion systems and internal combustion engines[4,5].
Autoignition of engine-relevant fuels, such as DME, at
practical conditions (i.e. high pressures and intermediate to
low temperatures) is of fundamental importance to emerging
technologies such as Homogeneous Charge Compression
FUEL PROCESSING TECHNOLOGY 89 (2008) 1244 –1254
⁎Corresponding author.Department of Mechanical and Aerospace Engineering, Princeton University, Princeton, New Jersey 08544-5263,
USA. Tel.: [phone]; fax: [phone].
E-mail address:[email protected](M. Chaos).
0378-3820/$–see front matter © 2008 Elsevier B.V. All rights reserved.
doi:10.1016/j.fuproc.2008.05.021
available at www.sciencedirect.com
www.elsevier.com/locate/fuproc

2. Ignition (HCCI)[6]. Furthermore, interest and advancements in
fuel-flexible gas turbine power generation systems that use
DME[7]require the knowledge of its ignition and burning
characteristics. DME shock tube ignition studies are available
in the literature[8–11]; however, with the exception of the
studies of Pfhal et al.[8]and Zinner and Petersen[11], few DME
ignition results are available at pressure and temperature
conditions of practical interest.
In kinetic research, several detailed chemical models for
low and high temperature DME oxidation [12–17]have been
developed and validated against multiple experimental
observations. Of particular importance to the present study
is the ability of available chemical kinetic models to accurately
reproduce DME autoignition properties at engine-like condi-
tions. DME displays the classical two-stage, negative tempera-
ture coefficient (NTC) ignition behavior similar to that
observed with linear alkanes[6,8,18,19]. This behavior stems
from low temperature reactions involving hydrocarbon radi-
cals and molecular oxygen[19]. Therefore, a comprehensive
detailed DME kinetic model for gas turbine and engine
applications should correctly predict these low-temperature
autoignition features.
Due to the lack of ignition studies at elevated pressures and
low-to-intermediate temperatures noted above, the focus of
this work is to further the understanding of DME autoignition
behavior under such conditions. Autoignition experiments are
conducted for DME/oxidizer mixtures in a rapid compression
machine (RCM) over a range of compressed pressures,
compressed temperatures, and equivalence ratios. This
experimental dataset is then used as a basis for validation
and refinement of recently developed kinetic models for DME
oxidation, with special emphasis on prediction of autoignition
characteristics.
2. Experimental
The RCM system used in the present investigation has been
described in detail previously[20,21]and only a brief overview
will be given here. The RCM consists of a driver piston, a
reactor piston, a hydraulic motion control chamber, and a
driving air tank. The driver cylinder has a bore of 12.7 cm and
the reactor cylinder bore is 5 cm. The machine is pneumati-
cally driven and hydraulically stopped. The machine allows
variations of stroke and clearance height. The reaction
chamber is equipped with sensing devices for measuring
pressure and temperature, gas inlet/outlet ports for preparing
the reactant mixture, and quartz windows for optical access.
Additionally, the machine incorporates an optimized creviced
piston head design to promote a homogeneous and adiabatic
zone at the core of the reaction chamber[20]. Homogeneous
reactant mixtures are prepared manometrically inside a
mixing tank equipped with a magnetic stirrer.
DME/O
2/N
2mixtures are studied over the temperature
range of 615–735K,pressurerangeof10 –20 bar, and
equivalence ratio range of 0.43–1.5.Table 1lists the composi-
tions of gas mixtures tested herein. Note that inTable 1,
equivalence ratio (ϕ) is changed by altering the mole fractions
of O
2and N2while keeping a constant mole fraction of DME.
Here, equivalence ratio is calculated byϕ=3X
DME/XO2, where
X
DMEandX
O2are the mole fractions of DME and O
2,
respectively. In certain studies, especially those dealing with
emissions[22],ϕmay not be a good measure of mixture
stoichiometry in comparing properties of oxygenated versus
non-oxygenated fuels. In these cases, the definition of an
“oxygen equivalence ratio,”ϕ
Ω[22,23], becomes relevant. For
the purposes of the present study, however, the use ofϕis
sufficient as no comparisons to other fuels are performed and
the main interest is in examining the temperature and
pressure dependence of DME autoignition. For oxygenated
hydrocarbon fuels (i.e. those containing H, C, and O)ϕ
Ωis
defined as the amount of oxygen atoms required to convert all
C and H atoms in the fuel/oxidizer mixture to CO
2and H
2O
divided by the amount of oxygen atoms present in the fuel/
oxidizer mixture. For the present DME mixtures,ϕ
Ω=7XDME/
(X
DME+2XO2)orϕ Ω=7ϕ/(6+ϕ). The oxygen equivalence ratio is
also listed inTable 1. Note that the two definitions vary by no
more than 9% over the conditions studied. In this RCM
investigation, DME (supplied by Fisher Scientific) is 99.5%
pure; O
2and N2gases (supplied by Praxair) are of ultra high
purity (99.993% and 99.999%, respectively). For a given mixture
composition with known initial temperature, the compressed
gas temperature at the end of the compression stroke (top
dead center, TDC),T
c, is varied by altering the compression
ratio; whereas the desired pressure at TDC is obtained by
varying the initial pressure of the reacting mixture for a given
compression ratio. The temperature at TDC is determined
from measured pressures by the adiabatic core hypothesis
according to the relation
Z
Tc
T0
g
gϕ1
dT
T
¼Ln
Pc
P0

whereP
0is the initial pressure,T 0is the initial temperature,γ
is the temperature-dependent specific heat ratio, andP
cis the
measured pressure at TDC.
3. Numerical modeling
Due to heat loss to the combustion chamber walls, compression
in an RCM is not truly adiabatic, and the pressure also decreases
during the post-compression period. Consequently, a numerical
model that accounts for the effect of heat loss is required to
correctly simulate the RCM experimental data. For a properly
designed creviced piston, such as the one used in the present
experiments, isentropic core compression can be assumed and
the effect of heat losses can be represented numerically by
comparing the computed and measured pressure traces [21].
Compression may also be non-ideal due to small amounts of
piston blow-by, but the method described below also accounts
for this effect as well. In the present study, non-adiabatic effects
are expressed through an adiabatic expansion by prescribing an
Table 1–Molar composition and stoichiometries of gas
mixtures tested
DME O
2 N2 ϕϕ Ω
1 7 27 0.43 0.47
1 4 30 0.75 0.78
1 2 32 1.50 1.40
1245FUEL PROCESSING TECHNOLOGY 89 (2008) 1244 –1254

3. effective volume as a function of time, during and after
compression. The volume expansion parameters are deter-
mined empirically from experiments using nonreactive mix-
tures with the same heat capacity, initial temperature, and
initial pressure as the reactive cases. Details of this procedure
can be found in[21]and[24]. The empirical effective volume
parameters used for the specific DME experiments discussed
herein are available in[25].
Numerical modeling of the experiments is performed using
the SENKIN code[26]coupled with CHEMKIN II[27]libraries.
The RCM is modeled as an adiabatic system with volume
specified as a function of time. The modeling begins from the
start of the compression stroke, and the importance of
modeling both compression and post-compression processes
will be further elaborated upon below. A recently developed
kinetic model for DME oxidation and pyrolysis[17]is used for
simulating the present experimental measurements. This
model consists of 55 species undergoing 290 reversible
reactions and has been validated for a wide range of physical
conditions, including flow reactor species-time history pro-
files at pressures up to 18 atm and temperatures near 1000 K
[14,15]as well as flow reactor reactivity profiles at 12.5 atm and
initial reaction temperatures from 550 K–900 K[15], jet-stirred
reactor species profiles up to 10 atm over the temperature
range of 550 K–1100 K[10,12], pyrolysis in a flow reactor up to
10 atm[14,17], shock tube ignition delay measurements at low
and high pressures[8–10], species profiles in low pressure and
atmospheric burner stabilized flames[28–30], and laminar
flame speeds up to 10 atm [31–33]. Updates to previous
modeling work[14,15]have been continued by Curran and
co-workers[16], and the most recent version of this model[34]
is also used to produce comparisons.
4. Experimental and numerical results
4.1. Ignition delay time definitions and modeling approach
Fig. 1shows a typical pressure trace for the ignition of a DME/
O
2/N
2mixture (1/4/30 by mole) at initial conditions of 297 K
and 591 Torr, along with the definitions of the first-stage
ignition delay (τ
1) and the overall ignition delay (τ). Specifi-
cally, the ignition delays (τ
1,τ) are defined as the time elapsed
from the end of the compression stroke (t= 0) to the instant of
maximum pressure gradient deduced from the pressure
history. For the case shown inFig. 1, the measured pressure
and the deduced temperature at the end of compression (t=0)
areP
c= 15.1 bar andT c= 651 K, respectively.
Fig. 2illustrates an example of the present RCM modeling.
Experimental and simulated pressure traces for both the
reactive mixture and the corresponding nonreactive mixture
with the same specific heat ratio are shown and compared in
Fig. 2. As explained above, based on the pressure trace of the
nonreactive experiment, the parameters required for the heat
transfer model (i.e. the effective volume parameters) are
deduced[21]. The experimental and computed pressure traces
for the nonreactive mixture are seen to match very well for
both compression and post-compression events, indicating
the adequacy of the present heat transfer model. The
empirically-determined parameters are then used for simu-
lating the corresponding reactive case. Furthermore, the
present heat loss modeling approach ensures that along
with the pressure history, the temperature history is also
correctly simulated[24]. Compared to the experimental data,
the simulated pressure trace obtained using the model of[17]
for the condition ofFig. 1over-predicts the first-stage and
overall ignition delays.
4.2. CSP analyses and compression stroke effects
It is noted fromFig. 1that while the compression time is
~30 ms, approximately 45% of the total pressure rise occurs in
the last 2 ms of the compression stroke. This compression
history is desirable in order to minimize the extent of chemical
reaction during compression, as it may affect subsequent
observations after compression ceases in two ways. First, if
chemical induction processes that control the initial produc-
tion of radicals begin during compression, radicals so gener-
ated may affect the observations after compression ceases.
Secondly, if significant chemical enthalpy changes occur dur-
ing compression, the temperature,T
c, immediately after com-
pression will differ from that determined from the measured
Fig. 1–Pressure traces illustrating the ignition characteristics
of DME and the definitions of the first-stage and overall
ignition delays. Molar composition: DME/O
2/N2= 1/4/30.
Fig. 2–Experimental and simulated pressure traces (using
the model of[17]) for the conditions shown inFig. 1. Both
reactive and nonreactive conditions are shown.
1246 FUEL PROCESSING TECHNOLOGY 89 (2008) 1244 –1254

4. P
cand the assumptions of isentropic, non-reactive compres-
sion with heat loss. If only small amounts of induction
chemistry occur, negligible mixture chemical enthalpy
changes will result, but overall ignition may still be affected
by the small radical concentrations produced during compres-
sion. It is also well known from motored engine studies[35]
that reactions may occur during the compression stroke
producing intermediates species and heat release both of
which may affect the observed chemical evolution and
autoignition behavior over the engine cycle.
In interpreting RCM observations, it has frequently been the
case that the conditions used as reference for simulations have
been the experimentally observed pressure and temperature
at the end of compression (e.g.[36]). Recently, some investiga-
tors[37,38]have referenced autoignition observations to
effective pressure and temperatures empirically derived from
measured data. It is implicitly assumed in such approaches
that chemical changes from the initial charge conditions that
might occur during compression do not affect the calculations
of autoignition delay parameters. Some RCM studies have also
reported considerable fuel consumption during the compres-
sion stroke[39], inferring substantial chemical enthalpy
changes. However, no analyses are available in the literature
that investigate the significance of the above two sources of
perturbations on observed ignition delays. Such analyses have
been performed here and are summarized below.
In the present study, a Computational Singular Perturbation
(CSP)[40]methodology recently demonstrated by Kazakov et al.
[41]is used. As developed[41], the methodology is applicable to
the analysis of systems that can be modeled as constant volume
processes (e.g. which has until recently[42]been a common
assumption in modeling shock tube ignition delay). To analyze
RCM data, the technique was further modified to accommodate
systems with time-changing volume, as described by Li et al.
[43]. Similar modifications can be made to accommodate
modeling of shock tube phenomena under known pressure
history[42]. The implementation of the CSP methodology used
in the present study is briefly outlined below. The chemical
kinetic reaction system may be represented by the following set
of ordinary differential equations
dz
dt
¼gzðÞ;
wherez=[T
~
y
1y2…yn–1yn]
T
is the state variable vector,T
~
the
normalized temperature,y
ithe species mass fractions (ntotal),
andgthe overall reaction rate vector. Unlike prior implementa-
tions of CSP, the inclusion of temperature as one of the CSP state
variables is essential for the direct analysis of thermokinetic
feedback so that factors controlling ignition and heat release can
be unambiguously determined. At any given timetthe reaction
rate vector can be differentiated (i.e.“perturbed”)withrespectto
time, dg/dt=Jgso that a local Jacobian matrix is defined,J=dg/
dz. One can, thus, perform the following decomposition onJ:
J¼VLV
Ω1
;
whereV=(v
1v2…vnvn+1is the matrix of eigenvectors andΛthe
diagonal matrix containing eigenvalues. The differentiation
essentially yields a system of linear ordinary differential
equations forg. Using the decomposition above, the reaction
rate vector can be represented as a sum of individual modes:
gtþDtðÞ c
X
nþ1
i¼1
f
iv
iexpk
iDtðÞ;
wheref
iis the mode amplitude (indicating the mode impor-
tance) andλ
ithe corresponding eigenvalue (indicating the mode
time scale and physical behavior). It is evident from the
equation above that the sign of the real component of the
eigenvalues, Re(λ
i), provides information about the dynamics of
the system. The modes with negative Re(λ
i) are referred to as
stable (decaying) modes, while the modes with positive Re(λ
i)
are unstable (explosive) modes. The explosive modes control
the ignition behavior of the kinetic system[41].
Results from the CSP analyses applied to one of the most
reactive cases (i.e. highP
candT c) considered in the present
DME study are shown inFig. 3. As expected, two-stage ignition
behavior is clearly evident in the predicted temperature
Fig. 3–Temperature (dashed lines) and CSP eigenvalue
spectrum (solid lines) time evolution for ignition of a DME/O
2/
N
2mixture (1/4/30 molar composition) initially at 523 Torr
and 297 K (P
c= 20.1 bar,T c= 720 K). Real eigenvalues are
plotted corresponding to the leading modes (i.e. the modes
with highest amplitudes). The figure insert shows results
during the compression stroke.
Fig. 4–Modes with largest amplitudes obtained from CSP
analysis for the conditions shown inFig. 2at 0.5 ms before
the end of the compression stroke. Open bars correspond to
stable modes, solid bars to explosive modes. The value of the
real component of the associated eigenvalues is also shown.
1247FUEL PROCESSING TECHNOLOGY 89 (2008) 1244 –1254

5. profile. Also shown inFig. 3are the real parts of the
eigenvalues, Re(λ), corresponding to the leading modes (i.e.,
the modes with highest amplitudes). Of special importance
are the trends shown during the compression stroke (the inset
ofFig. 3). CSP analyses reveal that the dominant mode during
compression is a stable mode (i.e. Re(λ)b0), which quickly
loses its stable nature as the compression ends. Shown in
Fig. 4are the amplitudes of the leading modes at a time
corresponding to 0.5 ms before the end of the compression
stroke for the conditions shown inFig. 3. Note that while an
explosive mode with sufficiently small time scale (i.e. Re(λ)=
5.55 ms
−1
) is present, its amplitude is nearly five orders of
magnitude smaller than the leading stable mode and, thus,
can be considered negligible, even at this late stage during
compression. The explosive modes shown inFig. 4are mainly
due to initiation reactions involving fuel and molecular
oxygen.
CSP results shown inFig. 3after the compression stroke
reveal an interesting pattern and warrant further discussion.
DME ignition exhibits similar patterns to those observed by
Kazakov et al.[41]in studying the two-stage ignition of n-
heptane. In the beginning of the first stage, the ignition is
characterized by a single explosive mode. As the system
approaches the end of the first stage, in addition to the
existing dominant explosive mode, another, slower explosive
mode with a lower amplitude appears. The two explosive
modes collapse, with their Re(λ) exhibiting a rapid decrease
passing through zero. Hence, the two explosive modes lose
their explosive nature, indicating the end of the first stage.
After the end of the first stage, the system again develops a
single dominant explosive mode that controls the ignition
runaway during the second stage. The reactions governing the
explosive modes that define the first stage are all low
temperature branching processes involving the internal iso-
merization and decomposition of peroxy radicals, whereas the
second ignition stage is nearly exclusively driven by the
decomposition of hydrogen peroxide[41].
To further support the discussion above,Fig. 5shows fuel
and oxidizer profiles as well as the heat release rate during
compression and ignition under the conditions studied above
and shown inFig. 3(i.e.P
c= 20.1 bar,T
c= 720 K). It is evident
that, during the compression stroke, there is very little
chemical reaction and, consequently, no measurable exother-
mic/endothermic effects on system sensible enthalpy. How-
ever, during the compression stroke, radical initiation
processes do begin to occur that subsequently play a role in
the further development of the radical pool after compression
ceases. Thus, even though there is little overall chemical
reaction during compression, the ignition processes observed
thereafter may be responsive to chemistry occurring during
Fig. 5–Predicted (using the model of[17]) species profiles and
heat release during compression and ignition of DME/O
2/N2
mixture (1/4/30 molar composition) at compressed conditions
of 20.1 bar and 720 K.
Fig. 6–Effect of the compression stroke on modeled (using the
model of[17]) pressure traces for a DME/O
2/N2mixture (1/4/30
molar composition). Open symbols represent calculations
performed considering the RCM compression stroke; lines are
results obtained by initializing the calculations at TDC for the
compressed pressure and temperature conditions listed and
using the initial mixture composition. Heat loss effect is
included in both calculations.
1248 FUEL PROCESSING TECHNOLOGY 89 (2008) 1244 –1254

6. the compression stroke. Such chemistry effects are included
in the present modeling approach, as predictions are calcu-
lated based upon the initial conditions prior to compression
and the pressure history, corrected for heat loss, over the
entire experimental procedure.
To further elaborate on the effects of small radical
production during compression, model results are presented
inFig. 6using two different modeling constraints: 1) the entire
compression and post-compression processes are modeled
using the experimental initial mixture conditions; 2) only
post-compression processes are modeled using the thermo-
dynamic state at TDC (i.e.P
candT
c) as the initial pressure and
temperature, and the initial experimental reaction mixture
composition. While the effect of heat loss is accounted for in
both calculations, the two modeling predictions can differ
significantly in computed ignition delay following compres-
sion, especially for experimental conditions that result in
short ignition delays. For example, fromFig. 6, at compressed
conditions of 20.1 bar and 639 K, the two modeling approaches
yield ignition delays of 31.87 ms and 34.31 ms, respectively. At
720 K, however, these values are 2.34 ms and 3.69 ms,
respectively, almost a 60% difference.Fig. 7shows the time
histories of important radicals in the DME system (i.e. H, O,
OH, HO
2,CH
3) during and after the compression stroke for the
conditions shown inFigs. 3 and 5. It is noted that these
radicals are all present in sub-ppm (i.e. HO
2,CH
3) and sub-ppb
(i.e. H, O, OH) levels at the end of compression. The radical pool
developed during compression, however small, can have a
marked effect on characteristic ignition delay observations for
the pressures and temperatures studied here, simply because
at these conditions chemical induction processes have
characteristic times that are comparable to the observed
ignition delays.
While the radical pool formed in the compression stroke
has an important effect on the first-stage ignition delay, its
effect on the second-stage chemical activities is minimal.
Specifically, by defining the second-stage ignition delay,τ
2,as
the time elapsed from the end of the first-stage ignition to the
instant of hot ignition (i.e.τ
2=τ−τ 1), comparison of the
computed pressure histories shown inFig. 6using the two
above-mentioned modeling approaches shows that τ
2is
nearly independent of the chemical induction processes that
occur during compression.
The discussion above indicates that chemical induction
processes are influenced by even small perturbations to the
initial radical pool formed during compression. This result
supports the views of Dryer and Chaos[44,45]put forth in
analyzing the recently reported discrepancies between che-
mical kinetic predictions and experiments of syngas/air
ignition delay[46]. As mentioned above, some rapid compres-
sion studies have modeled their experiments as constant
volume processes using conditions at TDC[36]or experimen-
tally determined effective pressures and temperatures[37,38]
as initial reference parameters. These investigations[36,37]
also report small differences (within experimental uncer-
tainty) in computed ignition delays between these approaches
and ones that include compression and post-compression
(similar to the one used here.) However, the validity of
modeling rapid compression experiments in this manner
should always be verified since, as shown above, this can lead
to considerable differences, especially for mixtures with short
ignition delays, such as those studied by Mittal et al.[47].
Furthermore, recent shock tube studies [42,48]under
similar pressure and temperature conditions as those
studied here have drawn attention to the considerable pre-
ignition pressure rise that is observed in some shock tube
experiments. This pressure rise can be considered as an
adiabatic compression and treated in a very similar manner
as the rapid compression processes described above, leading
to substantial differences in the interpretation of the
measured ignition delays in comparison to the use of
constant volume (V)andinternalenergy( U) modeling
assumptions, an approach commonly adopted by most
shock tube ignition delay modelers over the last few
decades. This approach is reasonable in terms of ignition
delay interpretations at short residence times, but cannot be
applied in interpreting long ignition delay data or those
found for high energy density mixtures. Shock tube experi-
mentalists have contributed to fostering this latter mis-
application by continuing to apply the constant U,V
assumptions to such data, and in reporting much of the
shock tube ignition delay data without pressure history data
where pre-ignition pressure changes may be discerned. A
physical explanation of this behavior, however, has not been
givenintheserecentworks [42,48]. Michael and Sutherland
[49]have discussed at length how pre-ignition pressure
changes can arise from non-idealities due to boundary layer
interactions and residual velocities behind the reflected
shock at the end wall. It appears that at temperatures below
around 1000 K, where chemical induction processes are of
significance to measured ignition delays, these types of
perturbations result in significant departure of observations
from interpretation of the data as constant volume and
internal energy processes. Moreover, many observations of
ignition processes in this regime show them to be non-
homogenous in nature[45], leading to additional sources of
pressure rise. Further research is needed to fully document
the processes that lead to non-homogeneous ignition and
subsequent pre-ignition pressure rise in diatomic bath gases
and at high fuel/oxidizer concentrations.
Fig. 7–Evolution of main radicals (computed using the model
of[17]) during compression and ignition of a DME/O
2/N2
mixture (1/4/30 molar composition) at compressed conditions
of 20.1 bar and 720 K.
1249FUEL PROCESSING TECHNOLOGY 89 (2008) 1244 –1254

7. Therefore, validation of chemical kinetic models using
ignition delay data gathered in shock tubes and rapid
compression machines for undiluted mixtures under elevated
pressures and low-to-intermediate temperatures (similar to
those studied here) strongly depend on the proper interpreta-
tion of experimental data[44]. The above discussion also
supports that ignition delays measured under these condi-
tions cannot be directly compared between different RCM and/
or shock tube studies without accounting for these“facility
dependent”effects[48]. In summary, the present work shows
that it is important to simulate the entire processes occurring
in the experiment, including the compression stroke.
4.3. Experimental and modeling results
Figs. 8 and 9show the effect of equivalence ratio on measured
first-stage and overall ignition delays as a function of
compressed temperature for compressed pressures of 10 bar
and 15 bar, respectively. As mentioned earlier, equivalence
ratio is varied through the change in the mole fraction of O
2,
while keeping the mole fraction of DME constant. It is seen
fromFigs. 8 and 9thatτincreases for increasing equivalence
ratio at higher compressed temperatures, whereas at lower
compressed temperaturesτchanges little with variations in
equivalence ratio. It is also noted that the NTC behavior
becomes more pronounced under fuel rich conditions. This
behavior is different from ignition studies of fuel/air mixtures
Fig. 8–Effect of equivalence ratio on experimental first-stage
(open symbols) and total (filled symbols) ignition delays of
DME at a compressed pressure of 10 bar; circles—ϕ= 0.43;
triangles—ϕ= 0.75; squares—ϕ= 1.5.
Fig. 9–Effect of equivalence ratio on experimental first-stage
(open symbols) and total (filled symbols) ignition delays of
DME at a compressed pressure of 15 bar; circles—ϕ= 0.43;
triangles—ϕ= 0.75; squares—ϕ= 1.5.
Fig. 10–Effect of compressed pressure on experimental first-
stage (open symbols) and total (filled symbols) ignition delays
of DME for an equivalence ratio of 0.75; circles—P
c= 10 bar;
triangles—P
c= 15 bar; squares—P c= 20 bar.
Fig. 11–Measured (symbols) and computed (lines) ignition
delays for a compressed pressure of 10 bar as a function of
equivalence ratio. Simulations use the model of[17]. Dashed
lines correspond to open symbols. For comparison, results
obtained using the model of[34]forϕ= 0.75 are shown as gray
lines.
1250 FUEL PROCESSING TECHNOLOGY 89 (2008) 1244 –1254

8. (e.g.[8]) whereτin the NTC region decreases as the mixture
stoichiometry changes from lean to rich. This is due to the fact
that in the present study the fuel mole fraction does not vary
and NTC reactivity is determined by the amount of O
2
available in the system. It is well known[19]that the fuel
peroxy species formed in the NTC region dictate the amount of
radical branching. For the cases considered herein, rich
mixtures have longer ignition times (i.e. slower branching) in
the NTC region because peroxy radical formation is reduced
due to the lower amount of oxygen present in the system
compared to lean cases.
Ignition delay variations as a function of compressed
pressure are shown inFig. 10for an equivalence ratio of 0.75.
It is noted that the NTC behavior becomes more pronounced
at lower compressed pressures. Moreover, for the pressure
conditions investigated, similar to the trends shown inFigs. 8
and 9, the first-stage ignition delay appears to be independent
of pressure. Further analysis shows thatτ
1is also a weak
function of the mole fraction of oxidizer, leaving temperature
as the primary variable.
Further modeling results are compared with experimental
data inFigs. 11–13. The predictions are observed to be in good
agreement with the data and the general trends are properly
reproduced. That is, compared to the experimental data, the
model exhibits similar variations in first-stage and overall
ignition delays with varying compressed pressure and equiva-
lence ratio. For comparison with the above predictions, all
yielded using the model of Zhao et al.[17], representative
results using an updated model [34]based on previous
development[5,14–16]are also shown inFigs. 11 and 12for
ϕ= 0.75 and inFig. 13forP
c= 15 bar. Larger discrepancies are
seen between computed and measured ignition delays when
using this model, especially at the lower compressed
temperatures.
One of the trends observed inFigs. 11–13is that the model
of Zhao et al.[17], overall, reproduces the overall ignition
delays better, whereas the first-stage ignition delays are
consistently over-predicted. To further investigate this issue,
a brute-force sensitivity analysis is performed to determine
what reactions are controlling the ignition delay at com-
pressed conditions of 15 bar and 650 K for an equivalence ratio
of 0.75. The ignition delay sensitivity to each reaction in the
model is computed by multiplying and dividing each reaction
rate,k, by a factor of two, and computing the ignition delay (i.e.
τ
ign_uporτ ign_dwnfor results obtained by adjusting rates
upwards or downwards, respectively). The sensitivity coeffi-
cient is calculated on a logarithmic basis, as it is usually
reported in the literature, namelyS=∂Ln(τ)/∂Ln(k) or, in this
case,S=Ln(τ
ign_up/τ
ign_dwn)/Ln(4). The results of such an
analysis are summarized inFig. 14.
Fig. 14showsthatatthepressureandtemperature
conditions studied herein, both first-stage and overall ignition
Fig. 12–Measured (symbols) and computed (lines) ignition
delays for a compressed pressure of 15 bar as a function of
equivalence ratio. Simulations use the model of[17]. Dashed
lines correspond to open symbols. For comparison, results
obtained using the model of[34]forϕ=0.75 are shown as gray
lines.
Fig. 13–Measured (symbols) and computed (lines) ignition
delays for an equivalence ratio of 0.75 as a function of
compressed pressure. Simulations use the model of[17].
Dashed lines correspond to open symbols. For comparison,
results obtained using the model of[34]are shown for
P
c= 15 bar as gray lines.
1251FUEL PROCESSING TECHNOLOGY 89 (2008) 1244 –1254

9. delays are most sensitive to the isomerization of the
methoxymethyl-peroxy radical, CH
3OCH
2O
2, forming a hydro-
peroxy-methoxymethyl radical CH
2OCH2O2H. CH2OCH2O2H
may undergo two routes: 1)β-scission releasing two formal-
dehyde molecules and a hydroxyl radical (i.e. CH
2OCH2O2H=
2CH
2O + OH); 2) reaction with molecular oxygen to form the
O
2CH2OCH2O2H radical, which consequently isomerizes and
decomposes releasing two OH radicals through the reactions
O
2CH2OCH2O2H=HO2CH2OCHO + OH and HO 2CH2OCHO =
OCH
2OCHO + OH. This latter sequence provides chain branch-
ing at low temperatures. As seen inFig. 14, theβ-scission of
CH
2OCH
2O
2H is the main reaction“opposing”ignition; as tem-
perature increases this step dominates leading to the negative
temperature coefficient (NTC) region as the reactivity of the
system decreases since only one reactive hydroxyl molecule
is released.
Rate expressions for the kinetic steps described above were
optimized in the study of Zhao et al.[17]and any significant
changes are likely to degrade agreement with other systems
against which the model was validated. One possible change,
however, is to the decomposition of hydroperoxymethyl
formate (HO
2CH
2OCHO). Ignition delay times are most sensi-
tive to this reaction at low temperatures (Tb650 K, approxi-
mately), with essentially no sensitivity at higher
temperatures. This is also the case in other systems (e.g.
flow reactors, flames, etc.)[17]. In the study of Zhao et al.[17]
the rate of hydroperoxymethyl formate decomposition was
increased; however, based on present results this rate may
need to be increased further. An increase of a factor of two
would reduce computed first-stage ignition delays by approxi-
mately 40% while minimally affecting total ignition delays.
This change, however, would make the rate of hydroperox-
ymethyl formate decomposition considerably different from
the values recommended by Sahetchian et al. [50]for the
decomposition of organic hydroperoxides. Therefore, it is
noted that model predictions of present RCM data can be
improved by changing this reaction rate without affecting the
performance of model against other targets; however, further
theoretical and experimental studies of the hydroperoxy-
methyl formate decomposition reaction are needed to sub-
stantiate this change.
5. Conclusion
DME autoignition was studied in a rapid compression machine
under a wide range of stoichiometries as well as compressed
pressures and temperatures. Results show two-stage ignition
behavior and the presence of a NTC region that is accentuated
at lower pressures and for oxygen lean mixtures. First-stage
ignition delays were found to be insensitive to variations in
pressure and equivalence ratio. Available chemical kinetic
models were used to simulate experimental results. The
model of Zhao et al.[17], developed and validated a priori,
was found to provide good agreement overall; thus, extending
its range of validation. First-stage ignition delays were
consistently over-predicted, however. Analysis of present
RCM data coupled with results from model sensitivity
analyses have helped identify the decomposition of hydro-
peroxymethyl formate as a reaction that may need updating
to better reproduce ignition delays under the conditions of the
present study; further experimental and theoretical studies of
this reaction are needed.
Modeling the compression process accurately is important
to properly interpreting RCM ignition delay results. The use of
effective thermodynamic parameters when modeling RCM
ignition delay, as it is commonly done in the literature, cannot
account for the establishment of the initial radical pool during
compression and its effect on induction chemistry. This study
has demonstrated the absence of essential thermochemical
activity during compression; nonetheless, and for the first
time, it is shown that induction chemistry occurring during
compression, even resulting in negligible changes in the
chemical energy release of the mixture, can result in
substantive changes in the ignition delays observed after
compression. Hence the specific design of the compression
process in various RCMs as well as their heat loss character-
istics must be taken into account in comparing observations
among different experimental venues. This perspective also
applies to the interpretation of shock tube ignition delay
measurements at similar pressures and temperatures, i.e.
conditions where observations are sensitive to chemical
induction processes, which in general, are strongly sensitive
to experimental perturbations.
Acknowledgements
The authors would like to thank Prof. Henry Curran for
providing an electronic copy of the model used in this study
(Ref. 34) This work has been supported by the Air Force Office
of Scientific Research under Grant No. FA9550-07-1-0515 and
by the Chemical Sciences, Geosciences and Biosciences
Division, Office of Basic Energy Sciences, Office of Science, U.
S. Department of Energy under Grant No. DE-FG02-86ER13503.
Fig. 14–First-stage (s
1) and overall (s) ignition delay time
sensitivity to reaction rates for a DME/O
2/N2= 1/4/30 (ϕ=0.75)
mixture under the compressed conditions ofP
c= 15 bar and
T
c= 650 K.
1252 FUEL PROCESSING TECHNOLOGY 89 (2008) 1244 –1254

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