Closed-loop control of HCCI combustion for DME using external EGR and.pdf

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Closed-loop control of HCCI combustion for DME using external EGR and

rebreathed EGR to reduce pressure-rise rate with combustion-phasing

retard

Dongwon Jung a,⇑, Norimasa Iida b

a Graduate School of Science and Technology, Keio University, 3-14-1 Hiyoshi, Kohoku-ku, Yokohama, Kanagawa 223-8522, Japanb Faculty of Science and Technology, Keio University, 3-14-1 Hiyoshi, Kohoku-ku, Yokohama, Kanagawa 223-8522, Japan

h i g h l i g h t s

External and rebreathed EGR were useful to control combustion phasing in HCCI engine. Combustion-phasing retard by external EGR effectively reduces pressure rise rate. Benefits of two-stage ignition fuel for combustion-phasing retard were demonstrated. Stable stoichiometric HCCI operation could be achieved by closed-loop control.

a r t i c l e i n f o

Article history:

Received 7 April 2014Received in revised form 25 October 2014Accepted 28 October 2014

Keywords:

Homogeneous charge combustion ignition(HCCI)Closed-loop controlCycle-to-cycle variationsExhaust gas recirculation (EGR)Combustion-phasing retardDimethyl ether (DME)

a b s t r a c t

This study experimentally investigates the effects of the combustion phasing on the homogeneous chargecompression ignition (HCCI) combustion, and implements a closed-loop control of HCCI combustion toreduce pressure-rise rate (PRR) with combustion-phasing retard. The experiments were conducted usingdimethyl ether (DME) in a single-cylinder HCCI research engine equipped with an exhaust gas recircula-tion (EGR) loop for external EGR and a two-stage exhaust cam for rebreathed EGR. The results showthat a

maximum PRR (PRR max) and a maximum in-cylinder charge temperature decreases with combustion-phasing retard. However, excessive combustion-phasing retard leads to unacceptable coefficient of variation (COV) of CA50 and IMEP with partial-burn and/or misfire cycles. To dampen increasing cycle-to-cycle variations around the limit of combustion-phasing retard, the closed-loop control of HCCIcombustion was implemented using three feedback variables. Finally, stable stoichiometric HCCIoperation could be achieved with extensive combustion-phasing retard while maintaining acceptablePRR max with the higher level of IMEP.

2014 Elsevier Ltd. All rights reserved.

1. Introduction

The homogenous charge compression ignition (HCCI) combus-

tion is a controlled autoignition of the homogeneous mixturethrough compression. This combustion is capable of providing bothhigh efficiencies and very low emissions of nitrogen oxides (NOx)and particulate matter (PM)[1,2]. Therefore, HCCI engine is consid-ered as a high-efficiency alternative to spark ignition (SI) gasolineengine and as a low-emissions alternative to compression ignition(CI) diesel engine. However, several technical difficulties need to beovercome for the practical application of HCCI in productionengines. Among these, reducing pressure-rise rate (PRR) associated

with excessive heat-release rate at the high-load HCCI operationcontinues to be a major issue. HCCI combustion depends not onlyon the unique chemical-kinetic mechanisms of the fuel, but also on

the thermal condition that mixture goes through during compres-sion process. Due to the self-accelerating nature of chemical-kinetic rates as the temperature rises, HCCI combustion tendstoward nearly instantaneous heat release, which can result inunacceptable PRR. To reduce PRR, thus extending high-load HCCIoperation, previous studies have demonstrated the strong poten-tial of combustion-phasing retard [3–6]. The benefits of combus-tion-phasing retard could be realized when the combustion isretarded later into the expansion stroke, since the volume-expansion of the in-cylinder charge by piston motion increases.The greater volume and volume-expansion rate with combus-tion-phasing retard more strongly counteracts the temperature

http://dx.doi.org/10.1016/j.apenergy.2014.10.085

0306-2619/ 2014 Elsevier Ltd. All rights reserved.

⇑ Corresponding author.

E-mail address: [email protected] (D. Jung).

Applied Energy 138 (2015) 315–330

Contents lists available at ScienceDirect

Applied Energy

j o u r n a l h o m e p a g e : w w w . e l s e v i e r . c o m / l o c a t e / a p e n e r g y

http://dx.doi.org/10.1016/j.apenergy.2014.10.085mailto:[email protected]://dx.doi.org/10.1016/j.apenergy.2014.10.085http://www.sciencedirect.com/science/journal/03062619http://www.elsevier.com/locate/apenergyhttp://www.elsevier.com/locate/apenergyhttp://www.sciencedirect.com/science/journal/03062619http://dx.doi.org/10.1016/j.apenergy.2014.10.085mailto:[email protected]://dx.doi.org/10.1016/j.apenergy.2014.10.085http://-/?-http://-/?-http://-/?-http://-/?-http://-/?-http://-/?-http://crossmark.crossref.org/dialog/?doi=10.1016/j.apenergy.2014.10.085&domain=pdfhttp://-/?-


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rise associated with the early exothermic reactions of the hottestzone, causing the combustion of subsequently colder zones to bemuch more delayed than the hottest zone. This leads to a stagedcombustion event, thus reducing PRR. However, there are limitsto how far combustion-phasing retard can be exploited to reducePRR since too much combustion-phasing retard leads to theappearance of partial-burn and/or misfire cycles, resulting in unac-ceptable cycle-to-cycle variations of the combustion phasing and/or the indicated mean effective pressure (IMEP).

For extending high-load HCCI operation with combustion-phas-ing retard, a fast and accurate closed-loop control of HCCI combus-tion would certainly be required to precisely control thecombustion phasing as the fueling rate is varied [7]. Various tech-niques have been demonstrated for the use of closed-loop controlof HCCI combustion such as fast thermal management [8], variablecompression ratio [9], and variable valve timings [10]. Ultimately,all of these techniques adjust the in-cylinder charge compositionand/or temperature at intake valve closing (IVC) so that the in-cyl-inder charge autoignites at the desired crank angle. From this per-spective, it appears that exhaust gas recirculation (EGR) can beused beneficially to control the combustion phasing with changesin the fueling rate by implementing the closed-loop control of HCCIcombustion. Some form of EGR has been shown to help control thecombustion phasing and/or to provide other benefits for extendinghigh-load HCCI operation [11]. External EGR that is the exhaustgases recirculated to the intake charge tends to retard the combus-tion phasing: mainly due to (1) a presence of carbon dioxide (CO2)and water (H2O) that stand out with low ratio of specific heat(c = c p/c v), which reduces the compressed-charge temperature[12], and (2) a reduction of the intake oxygen (O2) concentrationassociated with the addition of external EGR suppresses the autoig-nition reactions [13]. Consequently, the combined effects of lowerc and reduced O2 concentration by external EGR addition primarilylead to combustion-phasing retard even at the high-load HCCIoperation [14].

On the other hand, internal EGR that is usually the residual

gases retained from the previous cycle tends to advance the com-bustion phasing mainly by means of its high temperature thatheats the incoming charge [15,16]. At higher loads, less internalEGR is required due to the higher residual gas temperature andthe increased in-cylinder surface temperature. However, the pres-ence of hot residual gas pockets retained from the previous cycle asa result of incomplete mixing has the potential to increase thethermal stratification, unless the potentially more intense mixinghomogenizes the core of the in-cylinder charge too quickly. Itshould be possible to broaden the naturally occurring thermalstratification by the use of negative valve overlap or exhaustrebreathing [17,18], and in this way realize the benefits of thermal

stratification for reducing PRR. Since combustion-phasing retardacts to amplify the benefits of a given thermal stratification, thepotential of combustion-phasing retard for reducing PRR increasesrapidly with increasing thermal stratification [4]. This suggeststhat a combination of extensive combustion-phasing retard byexternal EGR addition and enhanced thermal stratification byinternal EGR addition should have good potential to control thecombustion phasing by implementing the closed-loop control of HCCI combustion.

The main objective of this study is to experimentally investigatehow combustion-phasing retard influences the cycle-to-cycle vari-ations of the combustion phasing and IMEP using external EGR (referred to here as ‘‘EEGR’’) and rebreathed EGR (referred to hereas ‘‘REGR’’ – assuming the sum of the retained residuals from theprevious cycle and the exhaust rebreathing by two-stage exhaustcam) for implementing the closed-loop control of HCCI combus-tion. Following the description of the experimental setup, thispaper covers four areas:

1. The main causes leading to combustion-phasing retard areidentified, and discussed in detail.

2. The effects of combustion-phasing retard on HCCI combustionare demonstrated, and the reasons for the results from combus-tion-phasing retard are examined.

3. The behaviors of cycle-to-cycle variations for stable operationand unstable operation are compared, and the sources of cycle-to-cycle variations are explained.

4. Finally, based on above findings, the closed-loop control of HCCIcombustion is implemented as a possible method for a stablestoichiometric HCCI operation around the limit of combus-tion-phasing retard.

2. Experimental setup

2.1. HCCI research engine

Fig. 1 shows a schematic of HCCI research engine facility, and anillustration of the incoming gases during the intake stroke. Theengine used for this study is based on a single-cylinder, air-cooled,SI gasoline engine (HONDA GX340 K1), which has been applied toan electric generator. The engine specifications are listed inTable 1.In-cylinder charge pressure measurements were made with atransducer (AVL GM12D) mounted in the center of the cylinderhead in place of the spark plug. The pressure transducer signalwere digitized and recorded at 1 crank angle (CA) increments for64 consecutive cycles. The in-cylinder charge temperature is com-puted using the ideal gas law in combination with the in-cylindercharge pressure, the known cylinder volume, and the trapped mass

EGR Cooler

Two-Stage

Exhaust Cam

Injector

Intake Cam

External EGR Loop

Fuel Flow Meter

ThrottleB

EEGR REGRAir Fuelmfuel mair mEEGR mREGR

T air T EEGR T REGR

Air Flow Meter

Air Filter

C p,fuel C p,air C p,EEGR C p,REGR

ThrottleA

ThrottleD

ThrottleC

T intake

External EGR Flow Meter

(b)(a)

T EEGR

P intake

T exhaust T intake

P exhaust

P air

F air

F fuel

F EEGR

P cylinder

Throttle

Air filter

Throttle

Throttle

Throttle

T air

Two-stage

exhaust cam

Fig. 1. (a) A schematic of the HCCI research engine facility, and (b) an illustration of the incoming gases during the intake stroke.

316 D. Jung, N. Iida / Applied Energy 138 (2015) 315–330


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of the in-cylinder gases. The temperature derived from the mea-sured pressure is inherently the mass-averaged temperature andnot the temperature of the adiabatic core. The maximum in-cylin-der charge temperatures reported in this study are extracted fromthese mass-averaged temperature traces. The temperature of air,EEGR and intake was measured with 1.0 mmdiameter type-K ther-mocouples mounted in the manifolds with thermocouple junctionsprotruding into the center of the flows. To allow a fast and precisemeasure of the in-cylinder REGR temperature to changes in thecycle-to-cycle variations [19,20], the temperature of exhaust andexhaust rebreathing was measured by / = 0.15 mm fine diametertype-K thermocouple mounted in the exhaust manifold close toexhaust valve (as indicated by the red point with annotation of T exhaust in Fig. 1a). A geometry of this / = 0.15 mm type-K thermo-couple is shown in Fig. 2a and b compares the response time for /= 0.15 mm and / = 1.00 mm type-K thermocouples at a gas veloc-ity of 100 m/s. All fueling was accomplished using a gas injectormounted in the intake manifold approximately 75 mm off theintake valve. In order to accurately determine the amount of fuelsupplied to injector, the digital mass flow meter (fuel flow meter(accuracy of ±1.0%)) was installed on the fuel line for controllingthe flow rate of fuel (F fuel). The flow rate of air (F air) and the flowrate of EEGR (F EEGR ) were metered by each laminar flow meter

(air flow meter (accuracy of ±0.1%) and EEGR flow meter (accuracyof ±1.0%)). Based on these flow rate data for the fuel and the air, afuel/air-equivalence ratio (/) in this study is determined by calcu-lating the mass of the in-cylinder fuel and air at IVC (mfuel and mair).

The combustion phasing is mainly controlled by the mixturecomposition and the thermal state of in-cylinder charge at IVC inHCCI engines due to no direct control mechanism such as a sparkdischarge of SI engine or a fuel injection of CI engine [1,2]. Thus,

the mass and the temperature of in-cylinder gases at IVC need tobe considered separately in order to allow cause-and-effect rela-tionships to be clearly identified. For this reason, a mass-averagedtemperature of in-cylinder charge at IVC of nth cycle (T IVC) wasemployed in this study, which is calculated using Eq. (1) with aspecific heat at constant pressure (c p), where temperature has tobe absolute (K).

T IVC ðKÞ

¼mair c p;air T air þ mfuel c p;fuel T fuel þ mEEGR c p;EEGR T EEGR þ mREGR c p;REGR T REGR

mair c p;air þ mfuel c p;fuel þ mEEGR c p;EEGR þ mREGR c p;REGR

ð1Þ

mair: in-cylinder air mass at IVC (g).mfuel: in-cylinder fuel mass at IVC (g).mEEGR : in-cylinder EEGR mass at IVC (g).mREGR : in-cylinder REGR mass at IVC (g).T air: in-cylinder air temperature at IVC (K).T fuel: in-cylinder fuel temperature at IVC (K).T EEGR : in-cylinder EEGR temperature at IVC (K).T REGR : in-cylinder REGR temperature at IVC (K).c p,air: air specific heat at T air (J/g/K).

c p,fuel: DME specific heat at T fuel (J/g/K).c p,EGR : EGR specific heat (= c p,N2XN2 + c p,O2XO2 + c p,ArXAr +c p,CO2XCO2 + c p,H2OXH2O + c p,fuelXfuel).c p,EEGR : EGR specific heat at T EEGR (J/g/K).c p,REGR : EGR specific heat at T REGR (J/g/K).

Appendix A gives further details of the method used when set-ting T IVC of nth cycle during data acquisition, including an exampleof experimental results of n 1th cycle and nth cycle for firedoperation in Fig. A1.

The 50% burn point (CA50) was selected as an objective mea-sure of combustion phasing. CA50 was determined by integratingthe cumulative apparent heat-release rate (AHRR), assuming con-stant gas properties for each operating condition and disregardingheat transfer [21]. The AHRR was calculated from the in-cylindercharge pressure using the typical first law and perfect gas analysis.It should be noted that low-temperature heat release (LTHR) isexcluded when calculating AHRR for CA50, since it eliminates theneed to correct for heat transfer during the weak low-temperaturecombustion phasing which extends for many crank angles.Effectively, the reported combustion phasing refers to CA50 forhigh-temperature heat release (HTHR), starting at the crank angleof minimum heat-release rate between LTHR and HTHR. Presenting

Table 1

Engine specifications.

Parameter Value

Displacement (single-cylinder) 337.81 c m3

Bore 82 mmStroke 64 mmConnecting rod length 112 mmCompression ratio 10.76

Number of valves 2Maximum power 7.1 kW@3600 rpmMaximum torque 22.1 Nm@2500 rpm

= 0.15mm

63.2% at 589.6K

t 63.2%= 0.029s

t 63.2% = Response Time

Gas Velocity = 100m/s= 1.0mm

t 63.2%= 0.561s

(b)(a)

43

=0.15mm

M14

Exhaust Valve

= 27.2

= 31.0

Fig. 2. (a) Geometry of / = 0.15 mm type-K thermocouple mounted close to exhaust valve. (b) Comparison of the response time for / = 0.15 mm and / = 1.00 mm type-Kthermocouples.

D. Jung, N. Iida/ Applied Energy 138 (2015) 315–330 317

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CA50 referring to HTHR is considered more relevant and consistentfrom the standpoint of quantifying the combustion phasing of HCCIcombustion [11,14,22–25]. Because the AHRR is the differencebetween the heat released and losses due to heat transfer, it doesnot always provide an adequate estimate of the actual heat-releaserate (HRR). This is particularly important for conditions where theHRR is small relative to the heat-transfer losses, such as at thebeginning and end of combustion, especially at retarded combus-tion phasing. For this study, the heat release was computed in amore refined way from the ensemble-averaged pressure of 64cycles, using the Woschni correlation for heat transfer [21] (withcoefficients adjusted for HCCI operation) and accounting forchanges in heat capacity with the fuel/air-equivalence ratio (/)and the in-cylinder charge temperature. Using the ensemble-aver-aged pressure has benefits from the standpoint of noise on theheat-release trances. On the other hand, it can lead to overesti-mated burn duration if the cycle-to-cycle variations are large[3,5,22,24,25].

As also shown in Fig. 1, the engine is equipped with an EGR loopfor EEGR, which features an EGR cooler (water heat exchanger),and a two-stage exhaust cam for REGR. Fig. 3 illustrates valveopening duration and valve lift featuring a valve event for exhaustrebreathing by the two-stage exhaust cam. In addition to systemsfor EEGR and REGR, four butterfly valves were installed, which can

be used to fully or partially shut off the flow of the intake air(Throttle A), the intake (Throttle B), REGR (Throttle C) and EEGR (Throttle D). These four butterfly valves can precisely change themixture composition of in-cylinder charge at IVC. To serve as a fun-dament for the following sections, the data acquired for motoredoperation and fired operation are presented in Fig. 4 when adjust-ing the angle of Throttle A (HA) and Throttle D (HD) from 0 deg(fully close) to 90 deg (fully open) at 90 deg (fully open) of ThrottleB (HB) and Throttle C (HC). Fig. 4a shows the resulting changes inthe mass ratio of in-cylinder gases to charge at IVC for motoredoperation. It is clear that the mass ratio of EEGR significantlyincreases with increasing HD by replacing air and REGR over awhole range of HA. For fired operation, this increase of EEGR byopening Throttle D (increasing HD) leads to combustion-phasingretard, and eventually misfire, as shown in Fig. 4b. The maximumpressure-rise rate (PRR max) in this study is used as a measure of the limit of knock, and the knock is considered unacceptable whenPRR max > 0.4 MPa/CA. On the other hand, COV of IMEP is used as ameasure of misfire, and the misfire is considered when COV of IMEP > 9%. Note that the percentage basis standard deviation hasbeen corrected for the heat-transfer losses of a motored cycle, asexpressed by Eq. (2) [3,5,6,23].

COV of IMEP ð%Þ ¼ 100 Std: Dev: IMEP

ðIMEPfired IMEPmotoredÞ

ð2Þ

2.2. Fuel selection

HCCI experiments have shown that high-volatility (and gas-phase) fuels are beneficial to form a homogeneous mixture for con-ventional premixed HCCI even HCCI engine can operate with avariety of fuels, [1,2,11,14,23–25]. In addition, a recent study dem-onstrated two-stage ignition fuels, which exhibit the first heatrelease ‘‘LTHR’’ associated with cool-flame chemistry before themain heat release ‘‘HTHR’’, have an advantage for retarding com-

bustion phasing due to the higher temperature-rise rate just priorto the hot-ignition point by the presence of LTHR, compared to sin-gle-stage ignition fuels (e.g., iso-octane and gasoline) [5,6,11,15].With higher temperature-rise rates, more stable HCCI operationwith combustion-phasing retard could be expected. Table 2 liststhe property of DME compared to those of diesel fuel that repre-sents a typical two-stage ignition fuel [26]. It can be noted from

ExperimentMotored

1500rpm B=90deg C =90deg

Rebreathed EGR (REGR)

Open

External EGR

(EEGR)

Air

(a) (b)Open

O p en

Misfire

Knock

IMEPn≈0.38MPa

≈0.42MPa

≈0.44MPa

≈0.46MPa

≈0.48MPa

≈0.49MPa

≈0.50MPa

ExperimentFiredDME =1.0Qin=Variable1500 rpm

B=90deg C =90deg

Fig. 4. Effects of the angle of Throttle A (HA) and the angle of Throttle D (HD) on (a) the mass ratio of in-cylinder gases to charge at IVC for motored operation, and on (b) theoperating range for fired operation.

TDCBDC BDC TDCTDC

Intake

Exhaust Rebreathing

Exhaust

Rebreathing

Intake

-270 -180 -90 0360270180900

Crank Angle [oCA aTDC]

V a l v

e L i f t [ m m ]

Fig. 3. Illustration of valve opening duration and valve lift.



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Table 2 that DME has a lower combustion enthalpy than that of diesel fuel due to the oxygen content of the molecules, whichnecessitates a higher volume of DME to be injected in order to sup-ply the same amount of energy to that provided by diesel fuel. Thiscan pose a problem for practical application of DME, especiallywhen considering high-load HCCI operation. To obtain sufficient

energy over the load-speed range, it can be desirable to operatewith diesel fuel. However, for diesel fuel, elevated intake tempera-ture (135–205C) is required before significant vaporizationoccurs to minimize the accumulation of liquid fuel on surfaces inthe intake system, making it difficult to form the homogeneousmixture. Instead, the disadvantages of DME for low liquid densityand low lower heating value could be mitigated if the injectionduration is increased with advanced injection timing. For thesereasons, dimethyl ether (DME) was selected as the fuel in thisstudy due not only to fairly volatile (low boiling point), but alsoto its strong LTHR (high cetane number) [27,28].

2.3. Closed-loop control

A brief disturbance of the combustion phasing and/or T IVC inHCCI engines can lead to large changes in the combustion due tono direct control mechanism such as a spark discharge of SI engineor a fuel injection of CI engine [1,2]. These changes then amplifythe initial disturbances, and make it impossible to allow stableHCCI operation with combustion-phasing retard. Thus, the fastand accurate closed-loop control would certainly be needed to pre-cisely control HCCI combustion [29,30]. For this reason, as Fig. 5shows, this study designs the closed-loop control system for feed-back control of HCCI combustion using LabVIEW program due toits flexible and scalable design. Since practical and reliable vari-ables are necessary for the fast response closed-loop control [31–33], one input variable of (fuel/air-equivalence ratio (/)) and twooutput variables (50% burn point (CA50) and heat-release effi-ciency (ghr)) were selected as the suitable variables for feedbackcontrol based on a recent study of the authors [34]. Here theheat-release efficiency used in this study is the ratio of the cumu-lative heat release during the period from the start of LTHR to theend of HTHR (Q hr) to the lower heating value of the fuel supplied inthe gas phase (Q LHV), as shown in Eq. (3).

heat-release efficiency ¼ ghr ¼ Q hrQ LHV

ð3Þ

As can be seen in Fig. 5, the variables are applied to two propor-tional-integral (PI) controllers (for / and CA50) and three Calcula-tors (for /, CA50 and heat-release efficiency), and computed byeach Calculator after acquiringthe data from the in-cylinder chargepressure sensor and the two mass flow meters (air flow meter andfuel flow meter). However, the way to calculate the variable is onlydifferent for heat-release efficiency. For the variables of / andCA50, they were calculated by averaging three consecutive cycles just before the corresponding cycle. On the other hand, only whenhaving heat-release efficiency lower than 65% in the previous cycle,heat-release efficiency is eligible for the variable.

For stoichiometric HCCI operation, the reference value of /(/ref ) is set to 1.0, and then / is mainly controlled by PI controllerof / that adjusts the angle of Throttle A (HA). In addition, the ref-erence value of CA50 (CA50ref ) was set to 14.0 CA aTDC, and thenCA50 is controlled by PI controller of CA50 that adjusts the angle of Throttle D (HD) for HCCI operation around the limit of combustion-phasing retard (which is CA50 = 14.1 CA aTDC in this study).Except for the reference values of / and CA50, the threshold valueof heat-release efficiency is set to 65% in Calculator for heat-releaseefficiency, and the heat-release efficiency is calculated fordetermining whether the heat-release efficiency can be used as

(b)(a)OpenClose OpenClose

Experiment

1500 rpm

DME

Avg.64cycles

Experiment

1500 rpm

DME

Avg.64cycles

Fig. 6. (a) CA50 and (b) 10–90% burn duration as a function of the angle of Throttle D ( HD).

Throttle A

Throttle D

Pressure Sensor

Engine

+ –

+ –

CA50 ref

ref

+PI Controller

CA50

PI Controller

Calculator

CA50

Calculator

Calculator

heat-release efficiency

–

65% > ηhr

Air Flow Meter

Fuel Flow Meter

Fig. 5. Schematic of the closed-loop control system for feedback control using threevariables of equivalence ratio (/), 50% burn point (CA50) and combustion efficiency(gcomb).

Table 2

Properties of DME and diesel fuel [26].

Property (unit) DME Diesel

Chemical structure C2H6O –Molar mass (g/mol) 46.0 170.0Carbon content (mass%) 52.2 86Hydrogen content (mass%) 13 14Oxygen content (mass%) 34.8 0

Liquid density (kg/m3) 667 831Cetane number >55 40–50Autoignition temperature (K) 508 523Stoichiometric A/F (–) 9.0 14.6Boiling point at 1 atm (K) 248.1 450–643Lower heating value (MJ/kg) 27.6 42.5Kinematic viscosity of liquid (cSt)


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the variable or not. Finally, using three feedback variables, theclosed-loop control system adjusts HA and HD simultaneously,ultimately changing the mixture composition of the in-cylindercharge at IVC.

3. Results and discussion

All data presented in this study were acquired at 1500 rpmunder naturally aspirated operation with a nearly constant fuelingrate of 346 ± 2 J/cycle (12 mg/cycle). This operating point is a rel-atively high load for HCCI research engineused in this study. As dis-cussed in conjunction with Fig. 4 in the description of theexperimental setup, Throttle D alone can significantly increaseEEGR by replacing air and REGR, thus leading to combustion-phas-ing retard. In order to more clearly identify the causes for combus-tion-phasingretard, this study onlyadjustsHD from 0 deg to38 deg(HD = 0–38 deg) while maintainingtheangleof other throttlescon-stant at 30 deg (partially open) for Throttle A (HA = 30 deg), and at90 deg (fully open) for Throttle B and Throttle C (HB = HC = 90 deg).Furthermore, to more clearly see the effects of HD on the in-cylin-der charge state at IVC, the definition of EGR rate used in this studyis the ratio of the sum of the mass of the in-cylinder EEGR (m

EEGR )

and REGR (mREGR ) at IVC to the in-cylinder charge mass at IVC(mcharge = mair + mfuel + mEEGR + mREGR ), as shown in Eq. (4).

EGR rate ¼ mEEGR þ mREGR ð Þ

mchargeð4Þ

It should be noted that as Fig. 4a shows, Throttle A directlyinfluences on mair, leading up to a significant change of /, espe-cially in the 0–38 deg range of Throttle D. Stoichiometric HCCI

combustion with EGR addition represents a natural upper fuelinglimit [35,36]. Furthermore, it can be desirable to operate with astoichiometric charge to allow the use of a three-way catalyst[37]. Thus, HA was kept at 30 deg to investigate whether it is pos-sible to combust the stoichiometric charge with acceptable PRR around the limit of combustion-phasing retard.

3.1. Causes for combustion-phasing retard

As mentioned above, the combustion phasing can be controlledby only adjusting HD. Fig. 6 shows how CA50 and the 10–90% burnduration (from 10% burn point for HTHR to 90% burn point forHTHR) change with HD. CA50 is retarded with opening ThrottleD, and becomes more sensitive to changes inHD as the combustionphasing is moved later into the expansion stroke. Also, it can benoted that the burn duration increases with HD, exhibiting verysimilar trend to that observed in CA50. Increase of the burn dura-tion would be expected to contribute to the steepening slope of thecurve for CA50 as CA50 is moved later. To analytically explain thebehavior of retarding CA50 with increasing burn duration, theresulting changes for the in-cylinder charge state at IVC are identi-

fied in Figs. 7 and 8. Fig. 7a plots the mass of the in-cylinder charge,air, fuel, EEGR and REGR at IVC (mcharge, mair, mfuel, mEEGR andmREGR ) as a function of HD, which corresponds to the data pointsin Fig. 6. A significant linear increase of mEEGR is observed alongwith opening Throttle D, while mair and mREGR decreases linearly.As evident in Fig. 7b, the EGR rate significantly increases withopening throttle D from 39.8% (HD = 0 deg) to64.7% (HD = 38 deg).Furthermore, / also increases with opening Throttle D, reachingthe fuel/air-equivalence ratio around / = 1 in the 34–38 deg range


Air (mair )EEGR (mEEGR )

REGR (mREGR )

Charge (mcharge)

Fuel (mfuel )

Experiment

1500 rpm

DME

Avg.64cycles

Experiment

1500 rpm

DME

Avg.64cycles

γ

Fig. 7. (a) Mass of the in-cylinder charge, air, fuel, EEGR and REGR at IVC (mcharge, mair, mfuel, mEEGR and mREGR ), and theresulting (b) EGR rate andequivalenceratio (/), whichcorresponds to the data points in Fig. 6.


Air (T air )EEGR (T EEGR )

REGR (T REGR )

Experiment

1500 rpm

DME

Avg.64cycles

Experiment

1500 rpm

DME

Avg.64cyclesT IVC

Intake (T fuel )

Fig. 8. (a) Temperature of the in-cylinder air, fuel, EEGR and REGR at IVC (T air, T fuel, T EEGR and T REGR ), and (b) mass-averaged temperature of in-cylinder charge (T IVC), whichcorresponds to the data points in Fig. 6.



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of Throttle D. This happens because the in-cylinder air mass is pro-gressively replaced with increasing the in-cylinder EEGR massunder the nearly constant fueling rate. Finally, it is noteworthy thatmcharge increases with opening Throttle D despite almost constantintake pressure. This happens because the large increase of the

in-cylinder EEGR mass raises the in-cylinder charge density, whichleads to observed increase of the in-cylinder charge mass for theconstant intake pressure. Fig. 8a plots the temperature of the in-cylinder air, intake, EEGR and REGR at IVC (T air, T fuel, T EEGR andT REGR ) as a function of HD, which corresponds to the data pointsin Fig. 6. As can be seen, T air, T fuel and T EEGR are almost constantin the 0–38 deg range of Throttle D. However, it should be notedthat

T REGR increases significantly with opening Throttle D in a

highly non-linear manner. Viewing changes in the mass and thetemperature of the in-cylinder gases with increasing HD make itpossible to explain the slope of the curve for T IVC in Fig. 8b. Com-paring changes of the mass and the temperature of the in-cylindergases at IVC shown in Figs. 7a and 8a, it can be concluded that thelarge increase of mEEGR with low T EEGR primarily reduces T IVC whilea non-linear increase of T REGR significantly contributes to a trend of decreasing slope of the curve for T IVC. Taken together, a combina-tion of significant reduction of T IVC andthe increased burn durationcoupled with increasing sensitivity of CA50 to changes in the com-bustion phasing during the expansion stroke leads to combustion-phasing retard. This can be clearly observed in Fig. 9, which showsthe in-cylinder charge pressure of 64 consecutive cycles forselected CA50 of 3.5, 1.6, 3.7, 7.7, 13.4 and 15.2 CA aTDC, cor-responding to HD of 0, 15, 25, 30, 34 and 36 deg in Fig. 6, respec-tively. As can be seen, combustion-phasing retard leads to theappearance of partial-burn and/or misfire cycles, and eventuallythe significant cycle-to-cycle variations of the in-cylinder chargepressure trace as the combustion phasing becomes overly retarded.

3.2. Results from combustion-phasing retard

The effects of combustion-phasing retard on HCCI combustionare examined in this section by considering the data points inFig. 6. Fig. 10a shows how maximum pressure-rise rate (PRR max)and maximum in-cylinder charge temperature (T max) changes withCA50. As can be seen, PRR max and T max decrease monotonicallywith combustion-phasing retard from the knock. This happens

mainly because when the combustion phasing is retarded, thegreater volume-expansion rate by piston motion more stronglycounteracts the temperature rise associated with the autoignitionof the hottest zone, which causes the autoignition of subsequentlycolder zones to be even more delayed than the hottest zone. Then,the combination of the greater volume-expansion of the in-cylin-der charge and the longer burn duration at later combustion phas-ing effectively reduce PRR max and T max. Also, EGR addition reducesthe compressed-charge temperature by lowing c and suppressesthe autoignition reactions by reducing O2 concentration, and thesecontribute as well to the reduction of PRR max and T max themselves.

I n - C y l i n d e r C h a r g e P r e s s u

r e [ M P a ]

CA50=-3.5oCA aTDC D=0deg

D=25deg

D=30deg

D=34deg

D=36deg CA50=15.2oCA aTDC

CA50=3.7oCA aTDC

CA50=7.7oCA aTDC

CA50=13.4oCA aTDC

CA50=-1.6oCA aTDC D=15deg

64cycles

64cycles

64cycles

64cycles

64cycles

64cycles

Average

Average

Average

Average

Average

Average

Fig. 9. In-cylinder charge pressure traces of 64 consecutive cycles for selected CA50of 3.5, 1.6, 3.7, 7.7, 13.4 and 15.2 CA aTDC, corresponding toHD of 0, 15, 25, 30,34 and 36deg, respectively.

MisfireKnock Operation MisfireKnock Operation

(b)(a)

PRRmax

Tmax IMEPg

IMEPn

PMEP

Fig. 10. (a) Maximum pressure-rise rate (PRR max) and maximum in-cylinder charge temperature (T max), and (b) Gross IMEP (IMEPg), net IMEP (IMEPn) and PMEP.


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This is an advantage for high-load HCCI operation. In addition toreducing both PRR max and T max, as Fig. 10b shows, gross IMEP(IMEP g ) and net IMEP (IMEPn) increase with combustion-phasingretard until CA50 of 7.7 CA aTDC, except for the earliest combus-tion phasing (CA50 of 3.5 CA aTDC) that shows the reduction of IMEP due to too large work transfer from the piston to the in-cyl-inder charge at the end of the combustion stroke. The curve forIMEP will be examined in greater detail later in this section. Onthe other hand, the pumping mean effective pressure (PMEP)decreases with combustion-phasing retard until partial-burn and/or misfire cycles start to appear, as plotted against the right-handaxis in Fig. 10b. This is thought to happen mainly because of areduction in the energy required to pump the charge through theengine by opening Throttle D, which would contributes to theincrease of mcharge under naturally aspirated operation, as shownin Fig. 7a. As Fig. 11 shows, CA50 and IMEP for 64 consecutivecycles are relatively stable until the limit of combustion-phasingretard despite the increase of the volume-expansion cooling withcombustion-phasing retard. Although COV of IMEP increasesslightly for earlier combustion phasing due to the increasedcycle-to-cycle variations of heat transfer caused by the cycle-to-cycle variations of knock intensity [3], COV of IMEP is only 1.21%at CA50 = 3.46 CA aTDC that is the middle of the combustion phas-ing range examined. From this perspective, it appears that combus-tion-phasing retard is an attractive way for extending high-loadHCCI operation. However, increasing cycle-to-cycle variations limithow far the combustion phasing can be retarded. The observedincrease of cycle-to-cycle variations with combustion-phasingretard shown in Fig. 11 can be logically explained. First, the

cycle-to-cycle variations of T IVC (presented by uncertainty bars inFig. 8b) primarily cause the cycle-to-cycle variations of the com-pressed-charge temperature. Even if there were no cycle-to-cyclevariations of T IVC, there would be the cycle-to-cycle variations of the compressed-charge temperature due to heat transfer and tur-bulent mixing during the compression stroke [38]. Furthermore,the cycle-to-cycle variations of the amount of trace species and/or unburned fuel recirculated from EGR add to the total cycle-to-cycle variations. These inevitable cycle-to-cycle variations will leadto variations in the autoignition timing and the burn duration fromcycle to cycle, and these variations will increase with combustion-phasing retard since temperature-rise rate becomes lower [3,6].Consequently, a certain cycle-to-cycle variations of the in-cylindercharge state (i.e. temperature, pressure, composition) will have alarger impact on the cycle-to-cycle variations of CA50 and IMEPat later combustion phasing. The sources of cycle-to-cycle varia-tions of CA50 and IMEP will be examined in detail in the followingsection by comparing the behaviors of cycle-to-cycle variations forstable operation and unstable operation.

To provide further insights regarding the effects of combustion-phasing retard by EGR addition, it becomes necessary to identifythe curve for IMEP in Fig. 10b more closely. This is done by exam-ining (1) the presence of strong LTHR for DME, and (2) thermal effi-ciency (gth) and heat-release efficiency (ghr). Examination on thepresence of strong LTHR is twofold. First, the presence of strongLTHR makes the amount of LTHR nearly constant even for delayedLTHR phase. Fig. 12 shows heat-release rate (HRR) for selected fivedata points, which corresponds to CA50 of 3.5 CA aTDC (knock),3.7 CA aTDC (minimum COV of IMEP), 7.7 CA aTDC (maximumIMEP), 13.4 CA aTDC (most retarded stable operation) and15.2 CA aTDC (unstable operation). As can be seen in Fig. 12a,the crank angle for onset of LTHR is somewhat delayed with com-bustion-phasing retard, but the amount of LTHR varies little withcombustion phasing, compared to that of HTHR. The net resulton LTHR is clear from Fig. 12b that shows a magnified view of HRR for LTHR presented in Fig. 12a. By carefully comparing LTHR

(Fig. 12b) and the corresponding in-cylinder charge pressure traces(Fig. 13a) in the 3010 CA aTDC range, it can be seen that thein-cylinder charge pressure during the compression stroke is sim-ilar until the onset of LTHR. Therefore, the more delayed LTHR phase with combustion-phasing retard leads to the higher in-cylin-der charge pressure while the in-cylinder charge temperaturepasses through the 720–890 K range where LTHR is most active[11,26]. Since the amount of LTHR increases with increasing pres-sure under high sensitivity to changes in pressure level, especiallyfor the fuels having strong LTHR, the temperature rise by LTHR becomes enough to make the hot-ignition point be reached even

CA50=3.7oCA aTDC

(Minimum COV of IMEP)

CA50=7.7oCA aTDC(Maximum IMEP)

CA50=15.2oCA aTDC

(Unstable)

CA50=13.4oCA aTDC

(Stable: Most retarded)

CA50=-3.5o

CA aTDC(Knock) 3.7oCA aTDC7.7oCA aTDC

15.2oCA aTDC

13.4oCA aTDC

CA50=-3.5o

CA aTDC

(b)(a)

Fig. 12. (a) Heat-release rate (HRR) for selected five data points of CA50 =3.5, 3.7, 7.7, 13.4 and 15.2 CA aTDC. (b) Magnified view of HRR for LTHR presented in Fig. 12a.

MisfireKnock Operation

CA50

IMEP

Fig. 11. COV of CA50 and COV of IMEP as a function of CA50.



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for delayed LTHR phase, which leads to the onset of HTHR. Second,the presence of strong LTHR increases the temperature-rise rate just prior to the hot-ignition point even at later combustion phas-

ing. Comparing HRR (Fig. 12a) and the corresponding in-cylindercharge temperature traces (Fig. 13b), it can be seen that the pres-ence of strong LTHR raises the in-cylinder charge temperature sig-nificantly, and leads to a relatively high temperature-rise rate justprior to the hot-ignition point. As mentioned above, a certaindegree of cycle-to-cycle variations of the in-cylinder charge statehas a different impact on the onset of HTHR depending on the levelof the temperature-rise rate just prior to the hot-ignition point.With higher temperature-rise rate, the timing of transition intohot ignition becomes less sensitive to the cycle-to-cycle variationseven at later combustion phasing [3,6]. Consequently, due to thepresence of strong LTHR for DME, a combination of the tempera-ture rise by nearly constant amount of LTHR even for delayed LTHR phase and the less sensitive of the onset of HTHR to the cycle-to-

cycle variations by high temperature-rise rate just prior to thehot-ignition point even at later combustion phasing contributesto the observed increase of IMEP with combustion-phasing retard.

Another important consideration for combustion-phasingretard by EGR addition in HCCI engines is its effects on thermalefficiency. Changes in thermal efficiency with EGR can result fromchanges in both combustion efficiency and thermodynamic effi-ciency, with changes in the thermodynamic efficiency resultingmainly from changes in combustion phasing, heat transfer, and c[24]. Here the thermal efficiency (gth) is defined in the classicalmanner with the definition of the combustion efficiency (gcomb)and thermodynamic efficiency (gtd) as:

thermal efficiency ¼ gth ¼ W

Q LHVð5Þ

combustion efficiency ¼ gcomb ¼Q enthalpy

Q LHVð6Þ

thermodynamic efficiency ¼ gtd ¼ W

Q enthalpyð7Þ

where W is the indicated work, and Q enthalpy is the estimated heatrelease based on the low heating value for the unburned CO andHC from the emissions measurements.

Note that:

gth ¼ gcomb gtd ð8Þ

To calculate Q enthalpy for the combustion efficiency and the ther-

modynamic efficiency, exhaust gases should be measured usingstandard exhaust-gas analysis equipment, but this laboratory does

not have the equipment to detect them directly. Therefore, theeffects of the combustion efficiency and the thermodynamic effi-ciency on the thermal efficiency will be explained by referring to

the results from previous studies. A comparison of the curve forIMEP in Fig. 10b and the curve for the thermal efficiency inFig. 14 shows that the general behavior of IMEP over the combus-tion phasing covered here is very similar to that of the thermal effi-ciency, as might be expected since the experiments in this studywere conducted under the nearly constant fueling rate (i.e. con-stant Q LHV). Two major observations can be made from the curvefor IMEP (or thermal efficiency), which increases, and thendecreases with combustion-phasing retard. First, IMEP increaseswith combustion-phasing retard. As discussed above inconjunction with Fig. 11, the changes in the knock intensity canhave a relatively large effect on heat transfer, thus affecting thethermodynamic efficiency. Although excessive PRR leads toincreased heat-transfer rates due to the flows associated with the

acoustic resonance of the knock, combustion-phasing retard canreduce the heat-transfer losses by reducing the PRR sufficientlyto stop knocking, thus increasing the thermal efficiency [3,23–25]. In addition, as discussed in previous studies [24,39,40], thecombustion efficiency generally increases with EGR addition sinceHC and CO recirculated from EGR get a second chance to combust.This can partly be explained by identifying the heat-release effi-ciency plotted against the right-hand axis in Fig. 14b. As can beseen, the heat-release efficiency increases with EGR addition (orwith combustion-phasing retard) until the limit of combustion-phasing, and this can be thought to contribute to the increase of the thermal efficiency with combustion-phasing retard. Second,

CA50=3.7oCA aTDC


CA50=7.7oCA aTDC

(Maximum IMEP)

CA50=15.2oCA aTDC

(Unstable)

CA50=13.4oCA aTDC


CA50=-3.5oCA aTDC

(Knock)

(b)(a)

CA50=3.7oCA aTDC


CA50=7.7oCA aTDC(Maximum IMEP)

CA50=15.2oCA aTDC(Unstable)

CA50=13.4oCA aTDC


CA50=-3.5oCA aTDC

(Knock)

Fig. 13. (a) In-cylinder charge pressure, and (b) in-cylinder charge temperature for selected five data points of CA50 =3.5, 3.7, 7.7, 13.4 and 15.2 CA aTDC.

MisfireKnock Operation

ηth

ηhr

Fig. 14. Thermal efficiency (gth) and heat-release efficiency (ghr) as a function of CA50.


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IMEP decreases with combustion-phasing retard. As mentionedabove, EGR addition tends to lower c mainly due to the presenceof CO2 and H2O. However, the lowered c has significant potentialfor reducing the thermodynamic efficiency [24,40]. Although thecombustion efficiency increases with EGR addition, this benefitfor the thermal efficiency is offset by the decrease of the thermody-namic efficiency caused by lowered c. Furthermore, the rate of increase of the combustion efficiency decreases with EGR additionsince the total amount of gases moving through the engine isreduced, which leads to a proportional reduction in CO and HCemissions. More importantly, combustion-phasing retard reducesthe effective expansion ratio and can potentially reduce the ther-modynamic efficiency. However, because the piston motion issmall around TDC, it is not until CA50 > 9.9 CAaTDC (HD = 32deg)that a substantial impact on the thermal efficiency is noted [23–25]. Typically, the combustion efficiency correlates with the com-bustion completeness, and T max on the order of 1500 K is neededto achieve a reasonable combustion efficiency with an accompany-ing complete of CO-to-CO2 reaction for IMEP [12,41]. Although acombination of lower T IVC by EGR addition and more volume-expansion cooling with combustion-phasing retard tend to reduceT max, all T max until the limit of combustion-phasing retard have acertain value above 1500 K, as shown in Fig. 10a. Based on this,the combustion efficiency only contributes to a small fraction of the decrease of the thermal efficiency. However, the combustionefficiency drops rapidly for T max below 1500 k, which leads toabrupt drop in IMEP as well as significant increase of COV of IMEP,as shown in Figs. 10b and 11.

3.3. Behavior of cycle-to-cycle variations for unstable operation

To better understand the behavior of cycle-to-cycle variationsfor unstable operation, this section compares the data set for 64consecutive cycles between two data points, which correspondsto CA50 = 13.4 CA aTDC (HD = 34 deg) representing the stableoperation case (COV of CA50 = 10.6%/COV of IMEP = 3.6%) and

CA50 = 15.2 CA aTDC (HD = 36 deg) representing the unstableoperation case (COV of CA50 = 21.2%/COV of IMEP = 23.2%). Thein-cylinder charge pressure traces for 64 consecutive cycles fromthese two data points can be identified and compared by the twographs presented below in Fig. 9. Fig. 15 first compares (a) CA50and (b) IMEP plotted against the cycle number for 64 consecutivecycles between stable operation and unstable operation. As canbe seen, stable operation has only small cycle-to-cycle variations,which are equally distributed to values higher and lower thanthe average. In contrast, unstable operation has much larger

cycle-to-cycle variations around the average while showing a cer-tain behavior of cycle-to-cycle variations. Careful examination of Fig. 15b shows that there are occasional partial-burn cycles withreduced IMEP, which are often followed by another more advancedpartial-burn cycle. Cycles #34–#41 and cycles #49–#53 are exam-ples of this. After several advanced partial-burn cycles, the mostadvanced partial-burn cycle are followed by another retarded cyclewith higher IMEP, and then being a well-burning cycle. However,comparing Fig. 15a and b, it can be noted that the cycle-to-cyclevariations of IMEP does not match well the cycle-to-cyclevariations of CA50 for unstable operation, especially for cycles#10–#33 that shows a nearly same level of IMEP despite excessivecombustion-phasing retard. This means that the cycle-to-cyclevariations of IMEP are insensitive to the cycle-to-cycle variationsof CA50. Therefore, the behavior of cycle-to-cycle variations of IMEP for unstable operation cannot be explained by only thebehavior of cycle-to-cycle variations of CA50. To more clearly showand compare the behavior of cycle-to-cycle variations of CA50 andIMEP in Fig. 15, CA50 and IMEP return map is plotted in Fig. 16 thatexhibits CA50 of next cycle against the CA50 (Fig. 16a) and IMEP of next cycle against the IMEP (Fig. 16b). As can be seen, CA50 andIMEP for 64 consecutive cycles of stable operation are presentwithin a fairly narrow CA50 and IMEP. This might not be too sur-prising due to the low COV of CA50 and IMEP with the highestheat-release efficiency. However, it should be noted that IMEP for64 consecutive cycles of unstable operation exhibits a certainbehavior of increasing cycle-to-cycle variations that turns clock-wise within a fairly wide IMEP range. The behavior of increasingcycle-to-cycle variations shown in Fig. 16 can be explained byexamining the cycle-to-cycle variations of the mass and the tem-perature of in-cylinder gases at IVC.

First, Fig. 17 compares (a) mair, (b) mEEGR and (c) mREGR plottedagainst the cycle number for 64 consecutive cycles between stableoperation and unstable operation. It is clear that the behavior of cycle-to-cycle variations of mair, mEEGR and mREGR has a same gen-eral trend to that of IMEP, although there are opposite behaviors

between IMEP, and mair and mEEGR . If there are the cycle-to-cyclevariations of IMEP with partial-burn and/or misfire cycles, thereshould be the cycle-to-cycle variations accompanying the increaseof mEEGR and/or the decrease of mREGR . However, the oppositebehaviors are observed in Fig. 17b and c. Therefore, the cycle-to-cycle variations of mEEGR and mREGR cannot explain why IMEPexhibits the certain behavior of increasing cycle-to-cycle variationsthat turns clockwise, as shown in Fig. 16b. Further examination onthe behavior of cycle-to-cycle variations of mair in Fig. 17a canprovide better insight into the behavior of cycle-to-cycle variations

(b)(a)

Unstable Operation (CA50=15.2oCA aTDC)

Stable Operation (CA50=13.4oCA aTDC)



Fig. 15. Comparison of (a) CA50 and (b) IMEP for 64 consecutive cycles between stable operation (CA50 = 13.4 CA aTDC) and unstable operation (CA50 = 15.2 CA aTDC).



11/16

of IMEP for unstable operation. As can be seen, mair increases withthe cycle-to-cycle variations of IMEP. This can lead to the reductionof / and the in-cylinder charge temperature, which favors the lar-ger cycle-to-cycle variations of IMEP. However, this also may notbe enough to explain the behavior of cycle-to-cycle variations of IMEP.

Second, Fig. 18 compares (a) T air, (b) T EEGR , (c) T REGR , and (d) T IVCplotted against the cycle number for 64 consecutive cycles



(b)(a)



53

36

37

38

39

41

35

52

51

4050

34

49

52

3751

38

39

5340

41

49

3650

35

34

No : Cycle Number No : Cycle Number

Fig. 16. Return map of (a) CA50 and (b) IMEP for 64 consecutive cycles, which corresponds to the data points in Fig. 15.

(a)

(b)

(c)



Fig. 17. Comparison of the mass of the in-cylinder (a) air, (b) EEGR and (c) REGR atIVC (mair, mEEGR and mREGR ) for 64 consecutive cycles between stable operation(CA50 = 13.4 CA aTDC) and unstable operation (CA50 = 15.2 CA aTDC).

(a)

(b)

(c)



(d)

Fig. 18. Comparison of the temperature of the in-cylinder (a) air, (b) EEGR and (c)REGR at IVC (T air, T EEGR and T REGR ), and (d) mass-averaged temperature of in-

cylinder charge (T IVC) for 64 consecutive cycles between stable operation(CA50 = 13.4 CA aTDC) and unstable operation (CA50 = 15.2CA aTDC).


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between stable operation and unstable operation. As mentionedbriefly in conjunction with Fig. 8a, T air and T EEGR are almost con-stant even for unstable operation. However, there are large cycle-to-cycle variations of T REGR showing a quiet similar behavior to thatof T IVC. This happens because T REGR significantly contributes to T IVCunder nearly constant T air and T EEGR . Therefore, the behavior of cycle-to-cycle variations of T IVC is thought to be occurred primarilywith a combination of the increase of

mair and

mREGR , and the

decrease of T REGR . As might be expected, the observed cycle-to-cycle variations of T IVC will typically lead to the cycle-to-cycle vari-ations of the onset of HTHR, especially when HTHR occurs laterinto the expansion stroke. So, the cycle following the partial-burnor misfire cycle has a higher probability for very late HTHR dueto the cycle-to-cycle variations of T IVC, likely resulting in a secondpartial-burn cycle. However, it should be noted that IMEP is recov-ered after several partial-burn cycles showing abrupt increase of IMEP at low T IVC for cycle #38 and cycle #52, as shown inFig. 16b. This happens because the presence of trace species and/

or unburned fuel is recirculated from one cycle to the next withthe REGR, and contributes to the autoignition enhancement, sincethe chemical effects would likely be much more prominent whenoperated with a drop of the combustion efficiency. Kaiser et al.[42] show that for HCCI combustion with low combustion effi-ciency, ethene, propene, iso-butene, and formaldehyde becomethe most prevalent exhaust hydrocarbon trace species, excludingfuel constituents. However, full analysis of the chemical effects of all the different trace species present in EGR is beyond the scopeof this study.

3.4. Closed-loop control of HCCI combustion

Based on the above findings, this section implements theclosed-loop control of HCCI combustion to dampen increasingcycle-to-cycle variations of CA50 and IMEP with combustion-phas-ing retard. As mentioned in the description of the closed-loop con-trol with Fig. 5, this study designed the closed-loop control system

Knock limit

Closed-Loop Control

Air (mair )

CA50 ref =14.0oCA aTDC

ηthreshold =65%

ref =1.0

EEGR (mEEGR )

Throttle A ( A)

Throttle D ( D

)

D=0deg 5deg 20deg 25deg 30deg 32deg 34deg 10deg 15deg 36deg

Fig. 19. Angle of Throttle A and Throttle D (HA andHD), mass of in-cylinder air and EEGR at IVC (mair and mEEGR ), CA50, fuel/air-equivalence ratio (/), heat-release efficiency

(ghr), maximum pressure-rise rate (PRR max), andnet IMEP (IMEPn) as a functionof the cycle number (#1–#640: without the closed-loop control/#641–#960: withthe closed-loop control).



13/16

for feedback control using three feedback variables of /, CA50 andheat-release efficiency for the fast response closed-loop control of HCCI combustion. Fig. 19 shows the results of 960 consecutivecycles for operation without (640 consecutive cycles from #1 to#640) and with (320 consecutive cycles from #641 to #960) theclosed-loop control of HCCI combustion. As discussed in conjunc-tion with Figs. 7 and 8, and can be clearly observed in Fig. 19 beforethe closed-loop control of HCCI combustion starts on cycle #641,

opening Throttle D while maintaining HA at 30 deg leads to thestair step appearance of the increase of mEEGR as well as thedecrease of mair, resulting in combustion-phasing retard. ExcessivePRR max around knock limit (=0.4 MPa/CA) at earlier combustionphasing decreases significantly with combustion-phasing retard,but IMEP is almost constant. However, since the cycle-to-cyclevariations become significantly larger with excessive combus-tion-phasing retard, the closed-loop control of HCCI combustionbegins to be needed to compensate for the cycle-to-cycle variationsaround the limit of combustion-phasing retard. After the closed-loop control of HCCI combustion starts on cycle #641, HA andHD are immediately adjusted by three variables (/, CA50, heat-release efficiency) based on the reference value of / (/ref = 1.0),CA50 (CA50ref = 14.0 CA aTDC) and threshold value of heat-release

efficiency (65% > ghr) during the following cycles. As can be seen,CA50 starts to wander back and forth around the reference valueof CA50ref = 14.0 CA aTDC, and the cycle-to-cycle variations of CA50 become less. The overall cycle-to-cycle variations of / seemsto be larger even after implementing the closed-loop control of HCCI combustion, but the observed variations are equally distrib-uted around the reference value of /ref = 1.0. In addition, theclosed-loop control of HCCI combustion allows most heat-releaseefficiency to be increased above 65%. This happens because open-ing Throttle A by the PI controller of / leads to the large increaseof mair by replacing mREGR , namely a substantial increase of O2concentration of the in-cylinder charge. Finally, to facilitate acomparison of the cycle-to-cycle variations of CA50 and IMEP afterimplementing the closed-loop control of HCCI combustion, the

results of these changes in CA50 and IMEP are replotted by CA50and IMEP return map in Fig. 20, which summarizes CA50 of next

cycle against the CA50 (Fig. 20a) and IMEP of next cycle againstthe IMEP(Fig. 20b). First, it can be observed that CA50 for 320 con-secutive cycles from cycle #641 to #960 with the closed-loop con-trol of HCCI combustion are present are present in a highly linearmanner within a relatively narrow CA50 range. As might beexpected from CA50 return map, COV of CA50 decreases from21.2% to 16.4% with more advanced average CA50 of 13.68 CAaTDC than the reference value of CA50 of 14.0 CA aTDC. Although

a combination of the cycles with advanced CA50 and the cycleswith mair increase may cause a higher PRR max, the significantlylower PRR max than knock limit can be achieved with the closed-loop control of HCCI combustion, and be maintained by the greatervolume-expansion rate of the in-cylinder charge and the longerburn duration at later combustion phasing. As expected based onthe reasoning above, COV of IMEP significantly decreases from23.2% to 4.7% by the closed-loop control of HCCI combustion whereCOV of IMEP above 9% is considered unacceptable. The net resulton IMEP is clear from IMEP return map inFig 20b, which comparesthe IMEP return map between operation without and with theclosed-loop control of HCCI combustion. It should be noted thatthe average IMEPn with the closed-loop control of HCCI combus-tion (=0.429 MPa) is higher than the maximum IMEPn of stable

operation without the closed-loop control of HCCI combustion(=0.408 MPa) corresponding to CA50 = 7.7CA in Fig. 10b. Conse-quently, the implementation of the closed-loop control of HCCIcombustion eventually makes it possible to achieve stable stoichi-ometric HCCI combustion around the limit of combustion-phasingretard while maintaining acceptable PRR with the higher level of IMEP.

4. Conclusion

This study experimentally investigates the effects of the com-bustion phasing on homogeneous charge compression ignition(HCCI) combustion, and implements a closed-loop control of HCCI

combustion to reduce pressure-rise rate (PRR) with combustion-phasing retard. The main conclusions of this study follow:

IMEP return mapStable

Operation

Unstable

Operation

Closed-Loop

Control

Symbol

Cycle Number [#]513~576

(64cycles)

577~640

(64cycles)

641~960

(320cycles)

COV of IMEP [%] 3.6 23.2 4.7

Average [MPa] 0.399 0.369 0.429

(b)(a)

CA50 return mapStable

Operation

Unstable

Operation

Closed-Loop

Control

Symbol

Cycle Number [#]513~576

(64cycles)

577~640

(64cycles)

641~960

(320cycles)

COV of CA50 [%] 10.6 21.2 16.4

Average [oCAaTDC] 13.4 15.2 13.68

Fig. 20. Summary graph showing the comparison of return map of (a) CA50 and (b) IMEP for stable operation without closed-loop control (64 consecutive cycles from #513to # 576) and unstable operation without closed-loop control (64 consecutive cycles from#577 to # 640), and the operationwith closed-loop control (320 consecutive cyclesfrom #641 to #960), which corresponds to CA50 and IMEP in Fig. 19. The data points for operation without the closed-loop control are reproduced from Fig. 16.


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1. Causes for combustion-phasing retard: The increase of the in-cylinder EEGR mass with low in-cylin-

der EEGR temperature at IVC primarily reduces the mass-averaged temperature of the in-cylinder charge at IVC (T IVC).

A non-linear temperature rise of the in-cylinder REGR at IVCsignificantly contributes to a trend of decreasing slope of T IVC curve.

A combination of significant reduction of T IVC

by EEGR addi-tion and the increased burn duration coupled with increas-ing sensitivity of 50% burn point (CA50) to changes in thecombustion phasing at later into the expansion stroke leadsto combustion-phasing retard.

2. Results from combustion-phasing retard: The combustion-phasing retard significantly reduces maxi-

mum pressure-rise rate (PRR max) and maximum in-cylindercharge temperature (T max), but also increases the cycle-to-cycle variations of CA50 and IMEP with partial-burn and/ormisfire cycles.

A combination of increased thermodynamic efficiency bylower PRR (i.e. reduced heat-transfer losses) and increasedcombustion efficiency by a second chance of CO and HCcombustion leads to the increase of IMEP with combus-tion-phasing retard.

A combination of decreased thermodynamic efficiency bylowered c and reduced effective expansion ratio, anddecreased combustion efficiency by lower T max than 1500 Kleads to the decrease of IMEP with combustion-phasingretard.

3. Behavior of cycle-to-cycle variations for unstable operation: Since the cycle-to-cycle variations of IMEP are insensitive to

the cycle-to-cycle variations of CA50, the behavior of cycle-to-cycle variations of IMEP does not match well the behaviorof cycle-to-cycle variations of CA50.

The behavior of cycle-to-cycle variations of IMEP is very sim-ilar to that of T IVC that increase with the increase of the massof the in-cylinder air and REGR at IVC, and the decrease of

the temperature of the in-cylinder EEGR and REGR at IVC. The IMEP is recovered after several partial-burn cycles at

low T IVC.4. Closed-loop control of HCCI combustion:

Stable stoichiometric HCCI operation around the limit of combustion-phasing retard can be achieved by implement-ing the closed-loop control of HCCI combustion.

Average IMEP for operation with the closed-loop control of HCCI combustion is higher than the maximum IMEP foroperation without the closed-loop control of HCCI combus-tion while maintaining acceptable PRR.

Stable stoichiometric HCCI operation around the limit of com-bustion-phasing retard can be achieved while maintaining accept-

able PRR with the higher level of IMEP by implementing theclosed-loop control of HCCI combustion. The control design for thisclosed-loop control systemhas low complexity compared to that of previous studies [7–10,30,31], which makes a fairly fast responsepossible to changes in desired / and CA50. Furthermore, theclosed-loop control using EGR can be broadly applied in other HCCIengines. This result provides the potential of the closed-loop con-trol of HCCI combustion using external EGR and internal EGR tooperate with a combination of the higher fueling rate and the moreextensive combustion-phasing retard for extending high-load HCCIoperation. However, to limit the parameter space, the experimen-tal investigation in this study has been conducted using DME,which is not common fuel, at a fixed engine speed of 1500 rpm.Both the fuel autoignitionreactivity and the engine speed influence

the observed cycle-to-cycle variations with combustion-phasingretard. Because of this, since the differences between the fuels

may grow or diminish for operation at other engine speeds, otherfuels must be evaluated before general conclusions can be drawn.This aspect should be considered when comparing the data pre-sented here with data acquired with other fuels at other operatingconditions in other engines.

Appendix A. Mass-averaged temperature of in-cylinder charge

at IVC (T IVC) of nth cycle

To demonstrate how T IVC of nth cycle was set during data acqui-sition, Fig. A1 shows an example of the experimental results of n-1th cycle and nth cycle for two consecutive fired operations. As canbe seen, for statistical treatment of T IVC of nth cycle, the calcula-tions related to the mass and the temperature of in-cylinder gasesat IVC of nth cycle were carried out over a 720 CA window fromIVC of n 1th cycle (hIVC(n1)) t o 1 CA before IVC of nth cycle (hIVC(-

n) 1). As Eqs. (A1)–(A3) show, integration of the flow rate of air,fuel and EEGR was performed over a 720 CA window to derivethe volumeof in-cylinder air, fuel and EEGR at IVC of nth cycle (V air,V fuel, V EEGR ). Since HCCI engine facility illustrated in Fig. 1 does nothave the equipment to directly measure REGR flow rate, the sum of

V air, V fuel and V EEGR (V air + V fuel + V EEGR ) then subtracts from the cyl-inder volume at IVC of nth cycle (V IVC) to approximate the in-cylin-der REGR volume at IVC of nth cycle (V REGR ), as expressed by Eq.(A4). With regard to the temperature of in-cylinder gases at IVCof nth cycle (T air, T fuel, T EEGR , T REGR ), the temperature of air, intakeand EEGR, and exhaust temperature were first measured over a720 CA window using each thermocouple, and then averaged. Itshould be noted that the average intake temperature and the aver-age exhaust temperature were assumed to be the in-cylinder fueltemperature at IVC of nth cycle (T fuel) and the in-cylinder REGR temperature at IVC of nth cycle (T REGR ), respectively. Then, themass of the in-cylinder gases at IVC of nth cycle (mair, mfuel, mEEGR ,mREGR ) is computed with the ideal gas equation of state for eachcombination of the volume of in-cylinder gases at IVC of nth cycle

0810810063 0-180360

Crank Angle [oCA aTDC]

-180 360

Intake

Exhaust Rebreathing Exhaust

Intake

Rebreathing

EEGR(F EEGR )

Air (F air )

θIVC(n-1)

∆ θ

= 720oCA

n-1th cycle nth cycle

Air

Exhaust

Fuel

(F fuel )

VolumePressure

θIVC(n)-1

Intake

EEGR

Fig. A1. Example of experimental results of n 1th cycle and nth cycle for fired

operation showing in-cylinder charge pressure and volume, the flow rate of air, fueland EEGR (F air, F fuel, F EEGR ), and the temperature of air, intake, EEGR and exhaust.



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(V air, V fuel, V EEGR , V REGR ), the temperature of in-cylinder gases at IVCof nth cycle (T air, T fuel, T EEGR , T REGR ) and the gas constant (Rair, Rfuel,REGR ), including the in-cylinder charge pressure at IVC of nth cycle(P IVC), as shown in Eqs. (A5)–(A8). This method used when settingT IVC of nth cycle during data acquisition can account for changes inT IVC that occur due to changes in the mass and the temperature of in-cylinder gases at IVC of nth cycle.

V air ¼I

1cycle

F airðhÞdh; R ¼ hIVCðn1Þ; hIVCðnÞ 1

ðA1Þ

V fuel ¼

I 1cycle

F fuelðhÞdh; R ¼ hIVCðn1Þ; hIVCðnÞ 1

ðA2Þ

V EEGR ¼

I 1cycle

F EEGR ðhÞdh; R ¼ hIVCðn1Þ; hIVCðnÞ 1

ðA3Þ

V REGR ¼ V IVC ðV air þ V fuel þ V EEGR Þ ðA4Þ

mair ¼P IVC V airRair T air

ðA5Þ

mfuel ¼P IVC V fuelRfuel T fuel

ðA6Þ

mEEGR ¼ P IVC V EEGR REGR T EEGR

ðA7Þ

mREGR ¼ P IVC V REGR REGR T REGR

ðA8Þ

P IVC: in-cylinder charge pressure at IVC of nth cycle (MPa).V IVC: in-cylinder charge volume at IVC of nth cycle (m

3).F air: air flow rate (L/s).F fuel: DME flow rate (L/s).

F EEGR : EEGR flow rate (L/s).V air: in-cylinder air volume at IVC of nth cycle (m

3).V fuel: in-cylinder DME volume at IVC of nth cycle (m

3).V EEGR : in-cylinder EEGR volume at IVC of nth cycle (m

3).V REGR : in-cylinder REGR volume at IVC of nth cycle (m

3).T air: in-cylinder air temperature at IVC of nth cycle (K).T fuel: in-cylinder fuel temperature at IVC of nth cycle (K).T EEGR : in-cylinder EEGR temperature at IVC of nth cycle (K).T REGR : in-cylinder REGR temperature at IVC of nth cycle (K).Rair: air gas constant (J/g/K).Rfuel: DME gas constant (J/g/K).REGR : EGR gas constant (J/g/K).mair: in-cylinder air mass at IVC of nth cycle (g).mfuel: in-cylinder DME mass at IVC of nth cycle (g).

mEEGR : in-cylinder EEGR mass at IVC of nth cycle (g).mREGR : in-cylinder REGR mass at IVC of nth cycle (g).

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Glossary

HCCI: homogeneous charge combustion ignitionDME: dimethyl ether

LTHR: low-temperature heat releaseHTHR: high-temperature heat releaseCA: crank angleaTDC: after top dead centerPI: proportional integralPRR: pressure-rise ratePRRmax: maximum pressure-rise rateT max: maximum in-cylinder charge temperature/: equivalence ratioCA50: 50% burn pointIVC: intake valve closingT IVC: mass-averaged temperature of the in-cylinder charge at intake valve closingIMEP: indicated mean effective pressurePMEP: pumping mean effective pressureEGR: exhaust gas recirculationREGR: rebreathed exhaust gas recirculationEEGR: external exhaust gas recirculation

COV: coefficient of variationghr: heat-release efficiencygth: thermal efficiencygcomb: combustion efficiencygtd: thermodynamic efficiencyc: ratio of specific heatc p: specific heat at constant pressurecv: specific heat at constant volume

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