Applied Thermal Engineering · 2013. 10. 8. · 3022 M. Zheng, S. Banerjee/Applied Thermal...

Applied Thermal Engineering 29 (2009) 3021–3035

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Applied Thermal Engineering

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Diesel oxidation catalyst and particulate filter modelingin active – Flow configurations

Ming Zheng *, Siddhartha BanerjeeMechanical, Automotive and Materials Engineering, University of Windsor, 401 Sunset Ave, Windsor, Ontario, Canada N9B 3P4

a r t i c l e i n f o a b s t r a c t

Article history:Received 22 April 2008Accepted 1 April 2009Available online 12 April 2009

Keywords:Diesel oxidation catalystDiesel particulate filterActive flow-control

1359-4311/$ - see front matter � 2009 Elsevier Ltd. Adoi:10.1016/j.applthermaleng.2009.04.017

* Corresponding author. Tel.: +1 519 253 3000x263E-mail address: [email protected] (M. Zheng).

A one-dimensional transient model is developed in order to carry out theoretical investigations on theactive flow diesel aftertreatment configurations. Simulations are carried out to predict the thermalresponse of particulate filters during active flow regeneration operations. Results indicate that the activeflow-control strategies can achieve higher energy efficiency in aftertreatment operations. The energy effi-ciency analysis is carried out using various active-flow configurations. The theoretical model is validatedusing the experimental results. Further empirical investigation is carried out in order to study energy effi-ciency of supplemental fuel in the active-flow configurations. Different engine operating modes are alsoinvestigated with the active-flow configurations. It is observed that diesel aftertreatment with active flowcan significantly improve in the supplemental energy efficiency.

� 2009 Elsevier Ltd. All rights reserved.

1. Introduction

In order to meet the stringent automotive emission regulations,increasing number of diesel vehicles around the world use catalyticconverters. Diesel exhaust aftertreatment applications include Die-sel Oxidation Catalyst (DOC), Diesel Particulate Filter (DPF) andNOx absorber. In order to enable aftertreatment operations, dieselengines commonly require supplemental energy to raise the other-wise relatively low exhaust temperature [1]. The appropriate after-treatment operation is largely dependent on the sufficienttemperature level of the substrate. In order to improve the overallenergy efficiency and keeping the most reactive conditions of theconverters of the aftertreatment operation systems, measures arebeing adopted such as the supplemental fuel delivery, the heatingof the exhaust gas and the active flow-control strategies that in-clude the periodic alternating exhaust flow paths through theaftertreatment device [1–4]. Depending on the vehicle applicationan effective and efficient active flow control scheme can be devel-oped with an objective to maintain conditions favorable to after-treatment operations.

The supplemental energy input can be efficiently applied in theexhaust stream by combining various active flow-control strate-gies such as parallel alternating flow, flow stagnation and periodicflow reversal to reduce the overall energy penalty drastically. Dif-ferent active flow schemes are shown in Fig. 1. It is important todevelop a suitable theoretical model that can simulate aftertreat-ment operations under these active flow control configurations.

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6; fax: +1 519 973 7007.

Several researchers have developed DOC and DPF models thatcan cater various flow control aftertreatment operation conditions[2,3]. The objective of the research presented in this paper is to de-velop a one-dimensional DOC–DPF model that can simulate after-treatment operations and behaviors under extreme conditions ofexhaust mass flow rates, oxygen concentrations and temperatures.In order to identify operating window for safe, effective and energyefficient regeneration, both numerical and experimental studiesare carried out. Further experiments are carried out in order to per-form an energy efficiency analysis of various active aftertreatmentoperations at low exhaust flow conditions.

The present study focuses on the evaluation of the energy effi-ciency on active-flow configurations of diesel aftertreatment sys-tems. The active flow schemes include partial flow control of theexhaust gas through the parallel or reversal flow aftertreatmentsystem with the addition of supplemental energy at the upstreamof the aftertreatment system, operating at various flow rates andexhaust gas temperatures. The theoretical and empirical resultsof aftertreatment behaviors under these conditions are presentedhere.

A parallel flow system with DOC–DPF units operates in alternat-ing cycles of filtration and regeneration with various space veloci-ties. An active regeneration with external supplemental energy isrelatively energy efficient with reduced space velocities, whenthe exhaust temperature is below the threshold limit for the after-treatment operations. By sharing the operating cycle with otherparallel devices, the partial flow restriction (PFR) operation caneffectively reduce the supplemental energy consumption to suc-cessfully carry out active regeneration. For energy efficient opera-tions, the PFR switches between different modes as shown in Table

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Nomenclature

C control volumeCDPF catalytic diesel particulate filterCFD computation fluid dynamicsCO carbon monoxidecp specific heat capacity at constant pressured hydraulic diameter of the channelD hydraulic diameter of DOC channelD1 inlet hydraulic diameter of DPF channelD2 outlet hydraulic diameter of DPF channelDOC diesel oxidation catalystDPF diesel particulate filterE activation energyFR flow reversalh convective heat transfer coefficientHC hydrocarbonka rate of adsorption coefficientks rate of reaction coefficientL length of the monolith channelLHV lower heating valueNc total number of cells in DOCNO nitrogen oxideNO2 nitrogen dioxideNOx oxides of nitrogenNu Nusselt numberO2 oxygenp pressurePDE partial differential equationsPFR partial flow restrictionPM particulate matterR ideal gas constantr reaction rate

RPM revolutions per minuteRFIR relative fuel injection rateSOF soluble organic fractionsSSF specific surface area of monolith as a fraction of total

volume of DOCSVF specific volume fractiont timeT temperatureTDMA tri diagonal matrix arrayTHC total hydrocarbon_min inlet gas space velocityq density of exhaust gasqgw density of exhaust gas in the filter walll viscosity coefficientk thermal conductivityuz exhaust gas velocity in the direction zvw exhaust gas velocity across the channel wallws channel wall thickness_waccum accumulated space velocity

Yi mole fraction of chemical specie denoted by I or otherspecies

Subscriptg gas phasei different chemical speciesj different reaction sitesp constant pressure states solid phasew filter wall1,2 inlet 1 and outlet 2 or thermodynamic state 1 and 2

3022 M. Zheng, S. Banerjee / Applied Thermal Engineering 29 (2009) 3021–3035

1.1. Different modes of flow control are made based on the require-ments of the regeneration. Generally mode 1 operation is stayedover longer period of time as the rate of filtration is slower thanthe rate of regeneration. The flow control valve is placed in orderto control the flow distribution during various modes of PFR oper-ation as shown in Fig. 2. The PFR operates in sequence from modes1 to 5.

Another active flow method is the periodic flow reversal opera-tion, which is, in essence, a heat energy trap, performs an activeheat recovery in addition to the heat retention capability of mono-lith solids. By cyclically alternating the direction of exhaust flow, athermal wave is produced at the frequency of flow reversal, so thatthe central substrate temperature retained regardless of the varia-tions in the exhaust gas temperature. The configuration of periodicflow reversal is shown in Fig. 3 [5].

EngineAir R

Engine tailors rawexhaust conditions satisfy aftertreatme

Aftertreatment modraw exhaust condit

with independent co

Raw ExhaustAirEngine

Engine cyclic λ controlExhaust temperature tempering

Overheat protection

Engine provides inputs to aftertreatment

controller

Fuel

(a). Passive Aftertreatment Control Process:

(b). Active Aftertreatment Control Process:

Fuel

M

Raw Exhaust

Fig. 1. Active vs. passive fl

2. Model formulations

This section presents the mathematical, chemical and physicalphenomenon that describes energy, mass and momentum ex-change characteristics in a single channel of the DOC and theDPF. The one-dimensional model equations are derived from themass, momentum and energy conservation equations of a singlechannel inside the monolith DOC and DPF. References are takenfrom the works of Triana, Johnson and Konstandopoulos [6,7]. Itis to be noted that over the past three decades, many researchershave developed one-dimensional transient models to simulatethe chemical- and thermo-physical responses of the aftertreatmentdevices. One of the original models was proposed by Bisset Edward[8]. He proposed basic one-dimensional model accounting the var-ious effects of flows inside the channels of DOC and DPF. Soon Kon-

Treated Exhaustaw Exhaust

to nt

ifies ions ntrol

Treated Exhaust

Active λ controlActive temperature control

Active flow control

PassiveAftertreatment

ActiveAftertreatment

Aftertreatment operation monitored by engine

controller

odified Exhaust

ow control strategies.

Table 1.1Operating modes of parallel flow.

Mode Injector Flow distribution Operation

Loop 1 Loop 2 Loop 1 Loop 2 Loop 1 Loop 2

1 OFF OFF 50 50 F F2 ON OFF 10 90 R F3 OFF OFF 0 100 S F4 OFF ON 90 10 F R5 OFF OFF 100 0 F S

F: filtration, R: regeneration, S: stagnation.

DPFDOCInjector 2

Tg in

Flow

Tg out

TreatedExhaust

ControlValve

Tg outTg in GasInlet

DPFDOCInjector 1

Exhaust

Fig. 2. Parallel flow configuration.

Exhaust gas in

Exhaust gas out

Forward Flow

Reverse Flow

4 way valve

Fig. 3. Flow reversal configuration.

Exhaust gas z

y

Washcoat

Substrate

Catalyst

Fig. 4. Single channel substrate.

M. Zheng, S. Banerjee / Applied Thermal Engineering 29 (2009) 3021–3035 3023

standopolous and his co-researchers developed the model pro-posed by Bisset and included chemical effects on the thermal re-sponse in the aftertreatment devices [9]. Other researchers suchas Johnson and Triana also contributed to the development ofDOC and DPF model over next decades [7,10–13]. Till date themathematical model developed to describe the thermo physicalbehavior in the DOC and DPF has been used to simulate aftertreat-ment operation under standard operating conditions. Compara-tively less literature is available on the model’s capability tosimulate active flow conditions such as flow stagnation that in-cludes extreme conditions of flow rate. This model is calibratedwith experimental data of thermal response of DOC and particulateregeneration of DPF.

2.1. DOC model

The Diesel Oxidation Catalysts are designed to chemically treatthe raw exhaust gas by the oxidation of pollutant species such ascarbon monoxide (CO), gas phase hydrocarbons (HC) and the vola-tile organic compounds such as soluble organic fraction (SOF) pres-ent in the diesel exhaust. Commonly used DOC is a honeycombmonolithic structure. The Nitrogen Oxide (NO) is also oxidized tonitrogen dioxide (NO2) at the temperatures from 270 �C to 430 �C[5]. The accumulated particulate matters (PM) are oxidized in thepresence of NO2 inside DPF that is commonly located at the down-stream of DOC in the aftertreatment systems. The effectiveness ofthe DOC in oxidizing CO and THC can be observed at temperaturesabove ‘‘light off” for the catalytic activity. At the temperatures be-low 250 �C, Palladium–Rhodium (Pd–Rh) catalysts that are com-monly used in DOC, show relatively low reactivity. At thetemperatures below ‘‘light off”, the surface reaction kinetics ofthe elemental chemical reactions governs the performance ofDOC, while as at temperatures above ‘‘light off”, the diffusion of ex-haust species from the gas to solid phase governs the performancecharacteristics of the DOC. The mass transfer process can be en-hanced by incorporating design changes to the DOC cell structure,by increasing the surface area for the mass diffusion between thegas and solid phases. The present study is conducted to simulatethe DOC performance in diesel exhaust condition with excessoxygen.

Three basic reactions are assumed to occur in DOC.

COþ 12

O2 ! CO2 ð1Þ

C3H6 þ92

O2 ! 3CO2 þ 3H2O ð2Þ

NOþ 12

O2 ! NO2 ð3Þ

The hydrocarbon oxidation is represented by the Eq. (2) wherethe exhaust hydrocarbon is represented by C3H6. Different rangesof hydrocarbon present in the diesel exhaust. Depending on thediesel combustion mode the range of hydrocarbon in the exhaustgas may vary. In order to model Diesel after-treatment, the exhausthydrocarbons may be assumed to mostly contain carbon chainsranging from C9 to C12 [14]. Therefore, considering the complica-tions in modeling HC reactions at the ranges present in the typicaldiesel exhaust, a lower molecular weight of hydrocarbon family isused to simplify the model calculations.

A one-dimensional numerical model has been developed tosimulate a single channel of the monolith substrate as shown inFig. 4. The primary objective of the model is to carry out the tran-sient thermal response simulation of the exhaust gas and the sub-strate wall. The model accounts for the convective heat transferbetween the solid and the gas phase, and the conductive heattransfer. The effects of thermal radiation from the substrate wallto the surroundings are not considered in the model due to itsintrinsic nature of simulation in one dimension along the directionof flow. In order to calculate the exothermic heat released from thechemical reactions, the chemical kinetics of the oxidation reactionsin the catalyst surface are considered.

Inlet ChannelRaw Exhaustgas in Accumulated Soot Layer1

Plug


The exhaust species balance in the solid phase is dependent onthe rate of mass diffusion, rate of catalyst surface adsorption andthe rate of disappearance due to chemical reaction. Assuming quasisteady state of the energy balance in the solid phase, the rate ofdisappearance of reactant species is equal to the rate of adsorptionof the species from the gaseous phase. The net rate of reaction istherefore controlled by the chemical kinetics when the rate ofadsorption is higher than the rate of chemical reaction – the condi-tion occurs mostly at the temperatures below light-off. However attemperatures above catalytic light-off, the rate of diffusion controlsthe rate of reaction since at these temperatures the rate of reactionis higher and thus the reaction is mostly controlled by the avail-ability of the reactants.

The model treats the exhaust gas as an ideal gas. The compress-ibility of the exhaust gas is not considered due to the low operatingpressures. The exhaust flow in the channel is assumed to be lami-nar due to low Reynolds number of the gas flow through the microchannels of the monolith substrate. In this paper, the effect of heatand mass transfer is simulated during heating of the substrate bythe exhaust gas at different levels of exhaust flow rate and flowreversal switching frequencies through the monolith. The gasphase mass balance and the energy balance between the gas andthe solid phase are derived from the mass and energy conservationlaws. The convective heat transfer rate between the solid and thegas phase is the most dominant form of heat transfer that isresponsible for the thermal behavior of substrate. In this particularstudy, the convective heat transfer coefficients is determinedassuming fully developed laminar flow of exhaust gas with con-stant surface temperature.

The mass conservation of the exhaust gas is written in Eq. (4) inthe partial differential form. The equation signifies that due to therise or fall in the temperature of the exhaust gas through the chan-nel, the exhaust gas density may vary. The variation in density willbe compensated by the velocity variation through the channel.

dðqguzÞdz

¼ 0 ð4Þ

The temperature variations of the exhaust gas through themonolith channel can be calculated from Eq. (5). The heat transferbetween the solid and the gas phase is primarily due to forced con-vection at the solid to gas interface.

SVFcp;g@ðqguzTgÞ

@zþ ðSSFÞhðTg � TsÞ ¼ 0 ð5Þ

The convective heat transfer coefficient is a function of the ex-haust gas conductivity, and the channel hydraulic diameter givenby Eq. (6). The Nu number is determined assuming fully developedlaminar flow of exhaust gas with constant surface temperaturewith the characteristic length of hydraulic diameter.

h ¼ Nukg

dð6Þ

Assuming a homogeneous solid medium of the substrate, fromthe Fourier’s Law of heat conduction, the partial differential equa-tion for the energy balance of the solid phase is derived and givenby Eq. (7).

ð1�SVFÞqscp;s@Ts

@t¼ksð1�SVFÞ@

2Ts

@z2 �hSSFðTs�TgÞþSSFX3

i¼1

ðDHiÞri

ð7Þ

Outlet Channel Filtered gas outFlow through wall

2

Fig. 5. Single channel DPF.

2.2. DPF model

The DPF design in various diesel applications is a ceramic por-ous wall flow honeycomb structure made of cordierite. The sus-

pended particulate matters are filtered due to the deep bedfiltration of the raw exhaust gas during its passage through theporous ceramic substrate wall. The unfiltered exhaust gas entersthe inlet channel of the DPF whose outer end is plugged. The ex-haust gas flows through the wall of the porous substrate resultingin the physical separation of the solid particulates of diameterexceeding 10 nm. The filtered exhaust gas is then allowed to passthrough the outlet. The schematic single channel layout of DPF isshown in Fig. 5.

This section describes the loading and regeneration behavior ofdiesel particulate filters using filtration of solid particulates andchemical kinetics of soot oxidation. The DPF model describes thesolid particulate matter deposition and subsequent oxidation ofsoot by O2 and NO2. The flow field calculation involves mass,momentum balance of the exhaust gas along the channel length.The initial boundary condition such as inlet pressure is calculatedby iteration technique till the values converge to satisfy the princi-ples of conservation for mass, momentum and energy.

The primary assumptions for the DPF model are similar to thosefor the DOC model except that the viscous friction is considered be-tween the exhaust gas and the channel wall. The chemical reac-tions in the DPF regeneration model are represented as a onestep global reaction given by Eqs. (8) and (9).

Cþ a1O2 ! 2 a1 �12

� �CO2 þ 2ð1� a1ÞCO ð8Þ

C þ a2NO2 ! a2NOþ ð2� a2ÞCOþ ða2 � 1ÞCO2 ð9Þ

The terms a1 and a2 are the coefficients of chemical selectivityfor the reactions of PM with O2 and NO2, respectively. The chemicalselectivity coefficients are dependent on the reaction rates ofrespective chemical reactions. In the Catalytic Diesel ParticulateFilter (CDPF), the reaction rate for the particulate oxidation withoxygen is enhanced and therefore the chemical selectivity of thecarbon oxidation with the available oxygen increases. The selectiv-ity coefficients are given by Eqs. (10)–(13).

a1 ¼ 1� fCO

2ð10Þ

a2 ¼ 2� gCO ð11Þ

fCO ¼1

1þ pf 1 � Yl1O2� eEf 1=RT

ð12Þ

gCO ¼1

1þ pf 2 � Yl2O2� eEf 2=RT

ð13Þ

where fCO and gCO are the selectivity coefficient of CO formationfrom the soot oxidation by O2 and NO2, respectively.

The mass, the momentum and the energy balance equations areseparately solved between the inlet and the outlet channels. Theentire DPF channel is discretised into numerous small control vol-umes to solve the partial differential equations (PDE) individuallyfor all the control volumes throughout the length of the DPFchannel.

The mass transfer of exhaust gas for the inlet and outlet chan-nels is given by Eqs. (14) and (15), respectively. The exhaust gas


density qgw in the filter wall is assumed to same as of correspond-ing exhaust gas density in the inlet channel as denoted by q1.

@

@zðq1u1Þ ¼ �

4D1

qgwvw ð14Þ

@

@zðq2u2Þ ¼

4D2

qgwvw ð15Þ

The momentum balance of the exhaust gas in the inlet and outletchannels of the DPF is given by Eqs. (16) and (17), respectively. It isassumed that the axial momentum is affected by the change invelocity magnitude and the friction between the exhaust gas andthe filter wall. The pressure drop across the filter wall between theinlet and the outlet channels is given by Darcy’s Law given by Eq.(18). The law states that the pressure drop across the filter wall isdependent on the wall flow velocity, the thickness of the wall con-sisting of the filter wall and the soot cake and their permeability.The dynamic viscosity of the exhaust gas is a temperature dependentvariable that contributes to the pressure drop across the wall. It is as-sumed that the exhaust gas viscosity is a function of the wall temper-ature. The solid temperature is assumed to be uniform throughoutthe thickness of the particulate cake and the filter wall.

@p1

@zþ @

@zðq1u2

1Þ ¼ �F

D21

lu1 ð16Þ

@p2

@zþ @

@zðq2u2

2Þ ¼ �F

D22

lu2 ð17Þ

p1 � p2 ¼ljp

vwwp þljs

vwws ð18Þ

The energy balance of the exhaust gas in the inlet and outletchannels is due to the net heat flux to the control volume, andthe heat transfer to the substrate wall by convection. The gas phaseheat transfers are given by Eqs. (19) and (20).

cp;gq1u1@T1

@z¼ 4

D1h1ðTs � T1Þ �

4D1

qgwvwcpwTs ð19Þ

cp;gq2u2@T2

@z¼ 4

D2h2ðTs � T2Þ þ

4D2

qgwvwcpwTs ð20Þ

The energy balance of the substrate wall is considered by takinginto account the convection heat transfer between the substrateand the exhaust gas at the inlet and the outlet channels, conduc-tion heat transfer in the solid phase and the heat generation dueto the exothermic soot oxidation. The solid phase in a loaded DPFchannel is composed of the substrate and the soot cake. But themodel assumes uniform solid temperature across the exhaust flowthroughout the thickness of the soot cake layer and the solid sub-strate layer within a control volume remains same. The effect ofconductive heat transfer across the soot layer is negligible to theeffects of conduction along the length of the DPF channel due tothe dimensional consideration.

@

@tððqwcpwws þ qpcppwpÞTsÞ ¼ h1ðT1 � TsÞ þ h2ðT2 � TsÞ

� kp@

@zwp

@Ts

@z

� �� ksws

@2Ts

@z2 þ _qs ð21Þ

With the mass, the axial momentum and the energy balanceequations for the inlet and outlet channels of DPF, the flow fieldcan be solved by numerical methods. The pressure drop is greatlyinfluenced by the filtration and the oxidation of the accumulatedsoot cake in the DPF inlet channel. Therefore it is important to cal-culate the soot cake thickness in order to determine the pressuredrop from the flow field calculation. The diffusion mass transferdominates the species transport in the porous wall and the sootcake [8]. The chemical reactions in the gas phase are neglecteddue to its short residence time inside the DPF channel. The molar

concentration of the reactant species like O2 and NO2 inside theDPF channel is assumed to be constant throughout the channellength for simplicity of calculation.

The conservation of the species can be expressed in terms of themolar concentrations of the reactant species and the rate of reac-tion in the embedded reaction sites. The total time derivative ofthe reactant species concentration is the sum of partial derivativeswith respective to time and distance across the flow (i.e. distancealong the particulate reaction site).

dðYkÞdt¼ @ð½Yk�Þ

@tþ @ð½Yk�vwÞ

@yk ¼ O2; NO2 ð22Þ

From Eq. (22), it is seen that the time rate of change of molarconcentration of reactants is governed by the balance betweenconvective transport across the wall (y direction) and the chemicalkinetics of the particulate oxidation. It is assumed that the chemi-cal reactions are limited to its kinetics due to the extremely smallreaction passage through the porous particulate filter layer. Exper-imental investigations by various researchers confirms that diffu-sion mass transfer rates are two orders of magnitude higher thanthe chemical reaction rates at the range of operating DPF regener-ation temperatures (450–1400 �C) [7].

Assuming the order of oxidation reaction as ak, the conservationequation of the reactant mole fractions are given by the steady stateconditions for the overall rate of species accumulation. Equating thetotal time derivative of the reactant species concentration to zero thereactant species concentration equation can be as follows.

@ðqgwYkÞ@t

¼ �@ðqgwYkvkÞ

@y� kk;jqgwakYk where kk;j ¼ Ak;jTseð�Ej=RTsÞ

ð23Þ

The local reaction rate of particulate oxidation is represented bythe modified Arrhenius form that is dependent on the temperatureof the reaction site and the activation energy. The activation energyEj varies between different sites of reactions such as the thermal,the catalytic and the soot cake, (the subscript j represents eachreaction site). It is assumed that the thermal reaction site isembedded inside the porous particulate filter wall.

Assuming quasi steady state of reaction, the transient term of theEq. (23) is neglected. Therefore the molar balance of the exhaust gasspecies across the DPF channel (y direction) is given by Eq. (24).

@

@yðqgwvwYkÞ ¼ �kk;jqgwYkak ð24Þ

The rate of O2 or NO2 depletion per unit wall surface area can beobtained by rearranging and integrating Eq. (24) across the wallthickness.

Rk ¼Z w

y¼0sjkk;jqgwYkdy ð25Þ

where sj is the specific surface area of soot particulate [13]. There-fore smaller the size of soot, greater will be the surface area of sootoxidation for a given mass of soot loading in a DPF channel. This willlead to a higher rate of oxidation with higher specific surface area ofoxidation.

The soot layer depletion rate during the regeneration process isa mathematical product of the rate of oxidation reaction and theorder of reaction. The overall rate of change of soot cake thicknessis given by:

qp@wp

@t¼ �

XK

MC

MK

� �RK þ _waccum ð26Þ

In the two layer model developed (Konstandopolous et al.,2000) the particulate layer thickness is divided into the thermallayer and catalytic layer [8]. A similar model treatment of the soot

Fig. 7. Schematic layout of simulation setup.


oxidation phenomenon is adopted to estimate soot oxidationembedded inside the porous filter walls. The filtration and theaccumulation mechanism of the exhaust particulate are modeledand integrated with the regeneration model in order to simulatesimultaneous regeneration and accumulation. The particulatesare initially deposited on the porous filter wall by Brownian diffu-sion through the direct interception that is considered the deepbed filtration. But as soon as the porous filter wall is saturated withthe accumulated soot, the particulate matter starts accumulatingover the filter wall to form soot cake inside the wall of the inletchannel. In the following simulations the accumulation mecha-nism is modeled based on the assumption that the soot depositionis dependent on the flow field of the exhaust gas at the inlet chan-nel and the filter wall. The rate of soot accumulation in a controlvolume is dependent on the influx of soot and the wall flow veloc-ity, and is inversely proportional to the inlet flow velocity. Higherthe inlet flow velocity, greater the chance of bulk soot been carriedover to the next control volume of the inlet channel.

u1@ð½Cg �Þ@z

¼ �kaccumð½Cg �Þvw ð27Þ

_waccum ¼ kaccumð½Cg �Þvw

u1

_min12 NC� � ð28Þ

The soot accumulation rate is estimated from the above equa-tion. The inlet exhaust mass is only passed through the inlet chan-nels of DPF which is only half of the total number of channels in theDPF (Nc). The mass flow rate of the exhaust gas at the intake of DPFgives an estimate of how much soot is carried over to DPF for thefiltration or the accumulation. In general high soot accumulationcoefficient (kaccum), represents the less flow ability of particulatesalong the inlet channel. The mass flow rate of the exhaust gas atthe inlet of DPF channel is given by:

_min ¼ qgu1ðt; 0ÞNC

2

� �ð29Þ

In the DPF model, the effect of porosity may change the pres-sure drop across DPF. However the combined porosity of the filterwall varies across the soot and porous filter layer. Variations in ex-haust soot property, however large, cannot affect the overall com-bined porosity, because the soot layer porosity is generallycontributes only 10–15% of overall porosity. Variations across thesoot layer and the filter wall are often neglected for thermal re-sponse calculation due to the complicacy of such model and alsodue to the fact that less than 1% of thermal variation may occuracross the soot layer and substrate wall. Therefore directional vari-ations are carefully considered for the model development asshown in Fig. 6.

The equations are solved numerically by applying computa-tional fluid dynamics (CFD) techniques using finite volume ap-proach. The PDE are discretised into simple arithmetic forms and

u1(z)Cb +z(1u)z( Δz)Cb(z+Δz)

Δz

vw(z)kaccumCb(z)

Porous Medium

Fig. 6. Soot accumulation in DPF filter wall.

the discretised equations are solved simultaneously within a timeloop of the simulation. The scalar variables such as temperature,pressure species concentration, etc., are solved along with vectorflow field parameters such as velocity and mass flux by semi-impli-cit pressure linked equation. However the energy equation for gasphase in DOC is solved using explicit method. The discretisedmomentum and energy equations are solved using tri-diagonalmatrix algorithm (TDMA) using forward elimination and backwardsubstitution for fast and efficient solution. The initial boundarycondition for pressure at the inlet of DPF is solved iteratively tillthe flow continuity convergence criterion is reached. The solutionat every iteration step is checked for convergences, boundednessand transportability to avoid solution divergence. In order to applythe pressure velocity equation in the DPF the channel is dividedinto a staggered mesh formation of small control volumes. The sca-lar variables are solved at every node whereas the vector variablesare solved at every faces using central differencing scheme. Thesimulation code setup in brief is given in Fig. 7.

3. Results

A single cylinder engine is set up with a DOC and a DPF to per-form tests on active and passive flow-control strategies to enableaftertreatment operation. Simulation results are compared withexperimental results in a DPF regeneration experiments performedin a dedicated DPF system as shown in Fig. 8. The experimental set-up is placed with thermocouples placed at various positions in or-der to record temperature along the length of DPF channel. Furtherexperiments and simulations reported in this work are carried outusing a DOC–DPF setup with the specification tabulated in Table3.1.

In order to validate the regeneration model, a simulation casewith initial soot loading of 5 g/L and initial substrate temperatureof 250 �C is compared with an experimental result. The initial sootload is assumed on the basis of the initial pressure drop for theDPF. Thermocouples are located at four different locations inside

Tg out

Thermocouple

Raw Exhaust

Exhaust Inlet

4DPF

3

2

1

Tg in

Treated Exhaust Gas

Fig. 8. Dedicated DPF setup.

Table 3.1DOC–DPF specification.

Value Unit

DOC diameter 5.66 in.DOC length 6 in.DOC wall thickness 8 milDOC cell density 400 cpsiDPF diameter 5.66 in.DPF length 6 in.DPF wall thickness 19 milDPF cell density 200 cpsiSubstrate material Cordierite –Catalyst alloy Pt–Pd –

0

5

10

15

20

25

30

35

50 100 150 200 250 300 350 400Exhaust Flowrate [g/s]

Pres

sure

Dro

p [k

Pa]

100 ˚C

200300

400

500

SimulationExperiment

Calibrating Pressure Drop6" by 5.66" 200 cpsi DPFSoot Loading: 0 g/LPressure drop with different inlet temperature

Fig. 9. Pressure drop vs. flow rate (0 g/L).

0

5

10

15

20

25

30

35

50 100 150 200 250 300 350 400Exhaust Flowrate [g/s]

Pres

sure

Dro

p [k

Pa]100 ˚C

200

300

400

500SimulationExperimentCalibrating Pressure Drop

6" by 5.66" 200 cpsi DPFSoot Loading: 5 g/LPressure drop with different inlet temperature

Fig. 10. Pressure drop vs. flow rate (5 g/L).

0

1

2

3

4

0 100 200 300 400 500Time [sec]

Pres

sure

dro

p [k

Pa]

0

25

50

75

100

Reg

ener

atio

n Ef

ficie

ncy

[%]

DPF initial temperature 250 °CInitial soot load 5 g/LTransient Flow Rate: 20~100 g/s

Experimental resultSimulation result

Fig. 11. Regeneration response comparison.


the substrate channel as given in Table 3.2. The model is validatedby tuning number of chemical kinetics, thermal and flow fieldparameters in order to match the pressure drops and the temper-atures of the substrate within the acceptable limits of tolerance.Regeneration model is validated by comparing the simulated ther-mal response and pressure drop of the DPF with the experimentaldata. Figs. 9 and 10 show that the simulation predicts the pressuredrop across DPF within ±5%.

From Figs. 11 and 12 it is seen that the simulation and theexperimental data are in reasonable agreement with one another.The simulated pressure drop is within ±0.25 kPa and the tempera-tures are comparable within ±70 �C with the experimental results.It is noticed that the timing of the regeneration is governed by thereaction kinetics. The simulation model is therefore able to predictthe timing of the regeneration response based on the chemicalkinetics calculations. Based on the validation some important cal-ibration parameters describing the thermo physical process duringDPF regeneration are tabulated in Table 3.3. Simulations are alsocarried with the same set of calibration parameters with lower ini-tial soot loading (3 g/L) and higher exhaust flow rate (�100 g/s).The initial pressure drop across DPF is measured 3.2 kPa as shownin Fig. 13. Simulation results with the following transient test arefairly accurate with the prediction of the pressure drop and regen-eration response in the DPF. The pressure drop across DPF initiallyrises because of the rise of exhaust temperature inside DPF chan-nel. The subsequent drop in pressure is due to the lower resistanceoffered to the wall flow due to depletion of soot cake thickness. TheDPF thermal response is shown in Fig. 14.

3.1. Simulation results

Simulations are carried out with the various active flow condi-tions for the DOC–DPF systems. The objectives of the theoretical

Table 3.2Thermocouple positions in DPF.

Solid temperature (�C) Location (z/L) in the monolith

Ts1 0.1Ts2 0.4Ts3 0.6Ts4 0.9

investigations are to carry out the thermal response and the energyefficiency analysis with different active flow strategies in the dieselaftertreatment system. The main focus of the simulation is to studyDPF regeneration and to get a better understanding of the variouspathways to initiate safe and efficient regenerations in the DPF.The energy required initiating such conditions in the DOC andthe DPF are quantified.

3.1.1. Gas temperature variationThe cases with different initial temperatures and the inlet gas

temperatures (Table 3.4) are simulated to analyze the energy re-quired to initiate and sustain regenerations inside DPF. The inlet

0

100

200

300

400

500

600

700

800

0 100 200 300 400 500

Time [sec]

Solid

tem

pera

ture

[o C]

Exhaust condition:DPF initial temperature 250 °CInitial soot load 5 g/L


Ts1Ts2Ts3Ts4

Fig. 12. Regeneration thermal response.

Table 3.3Calibration parameters.

Parameter Value Unit

F 28.45 –l 2.97 � 10�5 Pa sjs 1.54 � 10�15 m2

jp 5.55 � 10�15 m2

Nu 3.81 –cps 1.1 kJ/kg Kcpp 1.52 kJ/kg Kqs 1400 kg/m3

qp 140 kg/m3

ks 0.85 W/m Kkp 2.1 W/m Kkaccum 2500 –

0 100 200 300 400 5000

1

2

3

4

5

6

7

8

Time [sec]

Pres

sure

dro

p [k

Pa]

0

25

50

75

100

Reg

ener

atio

n Ef

ficie

ncy

[%]

Exhaust condition:DPF initial temperature 250 °CInitial soot load 3 g/L Experimental result

Simulation result

Fig. 13. Regeneration Response comparison.

100

200

300

400

500

600

700

0 100 200 300 400 500Time [sec]

Solid

tem

pera

ture

[o C

]

Exhaust condition:DPF initial temperature 250 ˚CInitial soot load 3 g/L


Ts1Ts2Ts3Ts4

Fig. 14. Regeneration thermal response.

Table 3.4Simulation conditions (Case S1).

Parameters Values UnitsCases

S1A S1B S1C

Exhaust gas temperature 310 300 290 �CInitial DOC temperature 310 300 290 �CInitial DPF temperature 310 300 290 �CInitial DPF soot load 5 g/LInlet THC 5000 ppmInlet CO 2000 ppmInlet O2 10 %Exhaust mass flow rate 10 g/s

0

200

400

600

800

1000

0 200 400 600 800 1000 1200Time [sec]

0

25

50

75

100

Reg

ener

atio

n Ef

ficie

ncy

[%]

Ts1 2 3 4

Soot Oxidation

T

DOC DPF

Ts1 Ts2Exhaust

Flow

T Ts3Ts4

Solid

tem

pera

ture

[oC

]

Fig. 15. Transient results (Case S1A).


temperature of the exhaust gas can be varied by installing an ex-haust gas heater similar to the one used in the experiment at theflow rates ranging from 2 to 15 g/s. The rise in the exhaust temper-ature depends on the exhaust gas flow rate and the energy supple-mented by the exhaust heater. The supplemental heat energy canbe effectively delivered at the low exhaust space velocities corre-sponding to the low flow rate loop of the PFR system. The experi-mental results indicate that if the DOC is placed before the DPF, thesetup will facilitate to accomplish the regeneration provided thatthe DOC attains catalytic ‘‘light off”. Once the ‘‘light off” conditionis reached, the DOC catalytic activity can be sustained by con-trolled addition of the supplemental fuel. The exothermic reactionsof fuel oxidation inside the DOC channels counterbalance the con-vective heat loss to the cooler exhaust gas. The exhaust gas picks

up the convective heat from the DOC channel and helps to sustainthe regeneration in the DPF in the downstream.

In the following simulations the DOC is positioned before theDPF so that the catalytic activity in the DOC helps to achieveregeneration conditions in DPF. A sufficient amount of externalfuel is injected in the upstream of DOC with the initial tempera-ture of DOC just above the temperature to sustain catalytic activ-ity. The simulation results of the transient thermal response alongwith the fraction of soot regeneration (i.e., regeneration effi-ciency) are presented for different cases. The soot profiles at dif-ferent time as the regeneration progresses are also shown fromFigs. 15–21. An energy efficiency comparison (Fig. 22) is madeby analyzing the fraction of soot regenerated with the total quan-tity of the supplemental energy delivered for the purpose. It is as-sumed that the supplemental fuel injected in the exhaust stream

0

200

400

600

800

1000

0 0.2 0.4 0.6 0.8 1

Normalized DPF channel length

Solid

tem

pera

ture

[oC

]

0 sec

100200 300

400500

Ts1 432

DOC DPF

Ts1 Ts2Exhaust

Flow

T Ts3Ts4

Fig. 16. Local wall temperature (Case S1A).

0

0.01

0.02

0.03

0.04

0.05

0 0.2 0.4 0.6 0.8 1


Soot

thic

knes

s [m

m] 0 sec

100

200

300400

500

Fig. 17. Soot oxidation (Case S1A).

0

200

400

600

800

1000

0 200 400 600 800 1000 1200

Time [sec]

Solid

tem

pera

ture

[o C

]

0

25

50

75

100

Reg

ener

atio

n Ef

ficie

ncy

[%]

Ts1 2 3 4

Soot Oxidation

T

DOC DPF

Ts1 Ts2Exhaust

Flow

T Ts3 Ts4

Fig. 18. Transient results (Case S1B).

0

0.01

0.02

0.03

0.04

0.05

0 0.2 0.4 0.6 0.8 1Normalized DPF channel length

Soot

thic

knes

s [m

m]

0 sec200

700

400

9001200

Fig. 19. Local wall temperature (Case S1B).

0 200 400 600 800 1000 1200

Solid

tem

pera

ture

[o C

]

0

200

400

600

800

1000

Time [sec]

0

25

50

75

100

Reg

ener

atio

n Ef

ficie

ncy

[%]

Ts1 2 3 4

Soot Oxidation

T

DOC DPF

Ts1 Ts2Exhaust

Flow

T Ts3 Ts4

Fig. 20. Soot oxidation (Case S1C).

0

0.01

0.02

0.03

0.04

0.05

0 0.2 0.4 0.6 0.8 1Normalized DPF channel length

Soot

thic

knes

s [m

m]

0 sec200

700

400

9001200

Fig. 21. Transient results (Case S1C).


is fully vaporized before entering to the DOC. The supplementalfuel delivery rate of 50 mg/s in an exhaust stream of approxi-mately 10 g/s corresponds to 5000 ppm by mass of THC by massfraction as indicated in Table 3.4. It is assumed that the supple-mental fuel is represented by the THC in the exhaust stream ofthe active flow control embedment.

The simulations are repeated with other inlet temperatures ofthe exhaust gas and the initial temperatures of the substrate. Theprogress of regeneration in DPF with the 5 g/L initial soot loadingis compared between the different cases of the initial substrate

temperatures and the inlet exhaust gas temperatures and shownin Fig. 22. It is observed that if the temperature is kept below theDOC ‘‘light off”, then the regeneration in DPF is not initiated. Thesimulation results demonstrate that the DOC–DPF system is capa-ble of the regeneration once the temperature is maintained above290 �C. The supplemental energy required to carry out the regener-ation towards the completion is given by Fig. 23. It is evident thatin order to initiate and complete the regeneration by energy effi-cient methods, an additional supplemental heater may be neededin the diesel exhaust system. The offset of additional heater cost,the PFR setup and the additional heat energy can be recoveredby the energy saved during the regeneration process in DPF.

0

20

40

60

80

100

0 200 400 600 800 1000 1200

Time [sec]

Reg

ener

atio

n Ef

ficie

ncy

[%]

Initial Temperature: 280 oC

290˚C

300˚C310˚C

Exhaust condition:Mass flow rate 10 g/sInitial soot load 5 g/LInlet THC 5000 ppm, CO 2000 ppm, O2 10 %

320˚C

330˚C

Fig. 22. Regeneration efficiency for different exhaust temperature.

0

0.5

1

1.5

2

2.5

3

0% 20% 40% 60% 80% 100%Regeneration Progress

Supp

lem

enta

l Ene

rgy

[MJ]

290˚C300˚C

310˚C 320˚C

330˚C

Variation of DOC inlet temperature:Flow Rate 10 g/sSupplemental Fuel 50 mg/sExhaust CO 2000 ppm, O2 10 %DPF initial soot load 5 g/LLHV of fuel = 42.7 kJ/g

Fig. 23. Energy efficiency analysis with regeneration progress.

0

0.01

0.02

0.03

0.04

0 0.2 0.4 0.6 0.8 1

Normalized DPF length

Wal

l flo

w v

eloc

ity [m

/s]

3006009001200

Time in secExhaust condition:Mass flow rate 80 g/sInlet PM 90 mg/m3

Fig. 24. Local wall flow (Case S2A).

0 0.2 0.4 0.6 0.8 10.00

0.02

0.04

0.06

0.08


Soot

thic

knes

s [m

m]

300

600

900

1200

Time [sec]Exhaust condition:Mass flow rate 80 g/sInlet PM 90 mg/m3

Fig. 25. Local pressure drop (Case S2A).


3.1.2. DPF soot accumulationThe operating condition of the exhaust and DPF is kept constant

while the exhaust gas flow rate and the soot concentration in theexhaust are varied as shown in Table 3.5. Between the cases S2Band S2C a comparison is made for wall flow rate and soot accumu-lation profile because between these conditions, the net soot filtra-tion are comparable. The conditions are summarized in Table 3.6.

Simulations are performed in order to generate soot accumula-tion profile under different conditions. The simulation model as-sumes that the soot accumulates over the existing soot cake (5 g/

Table 3.5Simulation matrix (Case S2).

Cases Soot

60 mg/m3 90 mg/m3

Flow rates 80 g/s S2B S2A53 g/s – S2C

Table 3.6Fixed exhaust condition (Case S2).

Parameters Values Units

Exhaust temperature 200 �CDPF initial temperature 200 �CExhaust O2 10 %Exhaust NO2 100 ppmInitial soot load 5 g/L

L initial soot load). The local wall flow velocity of the exhaust gasthrough the accumulated soot layer and the porous substrate bedis shown in Fig. 24 for Case S2A. The accumulated soot profile var-iation with time for Case S2A is presented in Fig. 25. The soot accu-mulation rates at different local positions (z/L ffi 0.01, 0.25, 0.5,0.75, 1) are compared for Case S2A and are presented in Fig. 26.It is been observed from the simulation results that the wall flowvelocity doesn’t vary with the progressive soot accumulation inthe inlet channel. This is consistent with the overall mass balanceof the exhaust gas. Because the overall mass flow rate of theexhaust gas is kept constant throughout the soot accumulationprocess, the pressure drop across the channel rise without the

0.000

0.005

0.010

0.015

0.020

0.025

0.030

0 200 400 600 800 1000

Time [sec]

Soot

acc

umul

atio

n ra

te [μ

m/s

]

0.0040.2520.50.7480.996

z/LExhaust condition:Mass flow rate 80 g/sInlet PM 90 mg/m3

Fig. 26. Local soot thickness (Case S2A).

0 0.2 0.4 0.6 0.8 10

2

4

6

8

10

12

14

Normalized DPF Length

Loca

l Pre

ssur

e D

rop

[kPa

]

3006009001200

Time [sec]

Fig. 27. Soot accumulation rate (Case S2A).

Table 3.7Simulation conditions (Cases S3).

Parameter Value Units

Cases S3A S3B S3C S3D

Exhaust mass flow rate 5 10 20 50 g/sExhaust gas temperature 300 �CInitial DOC temperature 300 �CInitial DPF temperature 300 �CInitial DPF soot load 5 g/LInlet THC 100 mg/sInlet CO 200 ppmInlet O2 10 %

0

1

2

3

4

5

0 0.2 0.4 0.6 0.8 1

Normalized DPF length

Wal

l flo

w v

eloc

ity [m

/s]

Exhaust Flow Rate 10 g/sInlet Fuelling 100 mg/s

0 sec

100 200

300

400

500

Fig. 29. Local wall flow (Case S3B).


change in the wall flow velocity. The local pressure drop across thechannel is dependent on the accumulated soot layer permeabilityand thickness. Therefore the pressure drop is increased with theaccumulation of soot in the inlet channel as shown in Fig. 27 forCase S2A. A comparison of the accumulated soot profile over a timeperiod of 1200 s is done and presented in Fig. 28.

3.1.3. Active flow DPF regenerationDPF soot regeneration is simulated using the same model in or-

der to calculate transient soot profiles and the related flow andtemperature field along the DPF channel under various active-flowconfigurations. In active-flow configurations, the exhaust gas iscontrolled to attain the suitable quantity and quality for an effi-cient and safe DPF regeneration. In order to initiate the regenera-tion, the external energy is supplemented to the exhaust streamin the form of fuel, electrical heating or microwave heating. Inthe following section, simulations are carried out by varying theexhaust flow rate and the temperature at the inlet of the aftertreat-ment in order to identify the regions sensitive to the regenerationthereby identifying the various energy efficient pathways to initi-ate DPF regeneration.

Keeping the amount of the external fuelling (supplementalmass rate) the same, variations in the exhaust flow rate result indifferent supplemental fuel fractions in the exhaust stream. Forexample, external fuelling of 100 mg/s is equivalent to 10,000and 2000 ppm by volume for exhaust gas flow rate of 10 and50 g/s, respectively, in the exhaust stream. In active flow controlschemes the exhaust gas flow rate and supplemental energy deliv-ery rate may be varied by changing parallel flow paths through theaftertreatment system. The simulation matrix is shown in Table

0.00

0.01

0.02

0.03

0.04

0.05

0.06

0.07

0.08

0 0.2 0.4 0.6 0.8 1

Normalized DPF Length

Soot

thic

knes

s [m

m]

Case S2B& Case S2CCase S2A

Fig. 28. Accumulated soot profile (1200 s).

3.7. A uniform soot-layer thickness at the initiation of the regener-ation is assumed as the initial boundary condition to the simula-tions presented in this section.

The progress of the DPF regeneration of the local soot layer andthe exhaust flow profile through the filter wall along the DPF chan-nel length are presented in this section. The local wall flow ratesand soot thickness profiles are presented in Fig. 29–31. The simu-lation cases are expanded to other exhaust mass flow rates keepingthe total supplemental fuelling rate constant. The simulation re-sults indicates that at higher flow rates (�50 g/s) the convectiveheat transfer from the cooler exhaust gas subside any surface cat-alyst activity in the DOC and subsequently resulting in lowerregeneration rate in the DPF placed downstream of the DOC. Theregeneration progress as a percentage of soot oxidized of the initialsoot load is presented for the various cases of exhaust flow rate inFig. 32. A subsequent energy analysis is carried out based on theregeneration progress and supplemental fuel delivered (assuming

0

0.01

0.02

0.03

0.04

0.05

0 0.2 0.4 0.6 0.8 1


Soot

thic

knes

s [m

m]

0 sec

100200

300

400

500


Fig. 30. Local soot thickness (Case S3B).

0

0.01

0.02

0.03

0.04

0.05

0 0.2 0.4 0.6 0.8 1


Soot

thic

knes

s [m

m]

0 sec

200

400

600

8001000


1200

Fig. 31. Local soot thickness (Case S3C).

0

20

40

60

80

100

0 200 400 600 800 1000 1200

Time [sec]

Reg

ener

atio

n Pr

ogre

ss [%

]

1020

5

25

1512

50

Exhaust mass flow rate [g/s] variation

Fig. 32. Regeneration progress (Case S3).

0

1

2

3

4

5

6

0 5 10 15 20 25 30Exhaust Flow Rate [g/s]

Ener

gy R

equi

red

[MJ]

20% 40% 60% 80% 100%Regeneration Progress

Fig. 34. Energy analysis (Case S3).


LHV 42.9 kJ/g) and is shown in Fig. 33. The analysis is also per-formed to quantify the energy required for the regeneration withthe variations in exhaust flow rates in Fig. 34.

From the wall flow and the soot profile transient simulation re-sults it can be observed that thicker soot layer offers more resis-tance to the passage of exhaust gas through the local wall region.This result is consistent with the observations made by variousother researchers [8,7]. It is also observed that with the increasein exhaust flow rate, soot regeneration is more uniformly distrib-uted along the channel length and conversely at lower flow ratethe regeneration occurs with local soot burning as shown in Figs.30 and 31. This can be explained that at lower exhaust flow rate,the oxygen supply to the regions of unburned soot is limited by

0

1

2

3

4

5

6

0 20 40 60 80 100

Regeneration Progress [%]

Supp

lem

enta

l Ene

rgy

Spen

t [M

J]

10

25 20 5

50

15

12

Exhaust mass flow rate [g/s] variation

Fig. 33. Energy for regeneration (Case S3).

the local exhaust flow through the soot layer and the filter wall.Other numerical investigations with the model reveal that at lowerflow rate, the soot regeneration is associated with local overheat-ing and significant loss of heat energy. On the other hand, at higherflow rates, the exothermic reactions in DOC are superseded by theconvective heat transfer affecting in loss of catalytic activity andsubsequently loss of regeneration conditions in DPF. From Fig. 34it is seen that there are specific conditions favorable to regenera-tion with the variations of exhaust flow rate. At high flow rates,due to low diesel exhaust temperature, the DOC catalyst doesnotachieve ‘‘light off” and at very low flow rates, the limited supplyof exhaust oxygen in the soot layer cannot provide favorable con-ditions to sustain regeneration.

Further simulations are carried out in order to investigate theinfluence of variations of the exhaust oxygen concentration (thatis subjected to the diesel engine load variation) on the regenerationefficiency during partial flow operation mode of active flowschemes. In this investigation, attempts are made to identify thecritical operating modes of partial flow system during transient en-gine operations. Due to minimal oxygen present in the exhauststream at high load engine conditions, aftertreatment reactionsin DOC and DPF may suffocate, thereby deactivating aftertreatmentoperation resulting in increased total energy consumption. On theother hand recovery of exhaust oxygen during transient high tolow load operation may also increase the chance of unsafe anduncontrolled regeneration in the DPF.

Simulation with different exhaust flow velocities and exhaustoxygen concentrations are carried within the conditions of partialflow operation. Supplemental energy to initiate regeneration (20%of regeneration) and subsequent completion of the regeneration(with 80% of initial soot been regenerated) are calculated andshown in Fig. 35. It is observed that exhaust oxygen concentration

0

0.5

1

1.5

2

2.5

3

3.5

0 5 10 15 20

Exhaust O2 [%]

Ener

gy R

equi

red

[MJ]

20% Regeneration80% Regeneration 15g/s

10 g/s

Exhaust Flow Rate

Exhaust Condition:Gas Temperature 300°CInitial DOC-DPF temperature 300°CInitial Soot Load 5 g/LExhaust CO 2000 ppmSupplemental Fuelling 100 mg/s

Regions deprived of O2

Fig. 35. Energy of regeneration.

188.8

377

125.5

251

0

100

200

300

400

500

10050Supplemental Fuelling [mg/s]

Exha

ust T

empe

ratu

re R

ise

[o C]

10 g/s15 g/s

Exhaust Flow Rate

Initial DOC Temperature 300 °CExhaust Temperature at DOC inlet 300°CExhaust O2 10 %

Fig. 36. Exhaust temperature in DOC.

0

1

2

3

4

0 100 200 300

Time [sec]

Rel

ativ

e fu

el in

ject

ion

rate

[%] Engine speed: 1400 RPM

Torque: 37 NmExhaust flow rate 9.9 g/sExternal fuel injection

Mode 1

Mode 2

Mode 3

Mode 3: ΔT = 284ºC Mode 2: ΔT = 287ºC

Mode 1: ΔT = 231ºC

Fig. 37. Modes of external fuel injection (1400 RPM).

400

500

600

700

mpe

ratu

re [

o C]

Engine speed: 1400 RPM Torque: 37 NmExhaust flow rate: 9.9 g/sExhaust O2: 11.6%

DOC 1

DPF 1DOC 2

DPF 2

DOC DPF1 2 1 2Exhaust Flow


plays a vital role in the supplemental energy efficiency in the con-ditions of partial flow operation due to limited availability of ex-haust oxygen for the regeneration reaction. Simulation alsoindicates substrate overheating (temperature above 1400 �C) with4 g/s and exhaust oxygen concentration of 20%. This indicates thereis a critical operating boundary for the DPF operation of duringload transients in partial flow restricting conditions.

In order to isolate the factors influenced by the exothermicreactions in DOC on the mechanism of DPF regeneration, the risein the exhaust gas temperature through the DOC channels withthe variations in the supplemental fuelling are calculated. It is ob-served if the DOC is operated above the catalytic ‘‘light off” condi-tions (i.e. �290 �C) the supplemental fuel is completely burntinside the DOC channels at the partial flow operating conditions.This is supported by the fact that at low flow rates, the exhaustgas has more time to burn completely in the DOC flow channelsthan normal passive flow conditions of high flow rates. Due to thisincrease in residence time of supplemental fuel in DOC, the rise inexhaust temperature is almost linearly proportional to the supple-mental fuelling rate, as shown in Fig. 36.

3.2. Experimental results

Empirical analysis is carried out in this study in order to furthervalidate on the supplemental fuel energy delivery strategies and itseffect in active-flow configurations. The thermal response of theDOC–DPF system is recorded at different supplemental energy fuelsupply rates at different engine speed and load conditions. Differ-ent modes of external fuel injection rate are applied to raise thetemperature of the DOC–DPF system to enable aftertreatmentoperation. A small gasoline injector, operated at 200 kPa pressure,is used for external supplemental energy delivery. An electric hea-ter (0.8 kW) heats up the exhaust stream to aid in diesel vaporiza-tion and mixing with the exhaust stream.

The space velocities of the exhaust gas across the DOC–DPF sys-tem at different engine operating conditions are given in Table 3.8.In a six-cylinder diesel engine these low space velocities will cor-respond to the conditions of partial flow restriction (PFR) acrossthe aftertreatment device. Therefore the experiments reported inthis paper are an appropriate representation of PFR.

Table 3.8Experimental conditions.

Engine speed (RPM) Engine torque (Nm) Space velocity (1/h)

1400 37 212801800 38 28373

Tests have been carried out to investigate the heating process ofthe DOC–DPF with three different external fuel injection rates attwo engine speeds. The engine operating conditions were keptsimilar at both the speeds to prevent any bias in the results. Thedifferent modes of external fuel injection at 1400 RPM are shownin Fig. 37. A relative fuel injection rate (RFIR) has been defined asthe ratio of the external fuel injection rate to the exhaust flow rate.These external fuel injection tests are carried out to attain a fixedtemperature rise (DT � 280 �C) in the DOC and to maintain the fi-nal temperature of the substrate at approximately 550 �C. Duringmodes 2 & 3, the fuel injection is initially higher to obtain fasterheating of the substrate. The fuelling rate is then reduced to thatof mode 1 to maintain the same final temperature for all themodes.

The overall thermal response of the substrate for the test isshown in Fig. 38. The system is allowed to cool down to the initialtemperature before starting the fuelling for the next testing mode.The results indicate that a higher fuel injection rate results in a fas-ter heating up of the substrate. On an energy efficiency basis, how-ever, the same may not hold true as shown later in the energyefficiency analysis.

The same tests were then repeated at 1800 RPM to investigatethe influence of the higher exhaust flow rate on the thermal re-sponse of the aftertreatment. The RFIR for the three modes ofexternal fuel injection is shown in Fig. 39 and the test results aregiven in Fig. 40. The temperature rise (DT � 340 �C) in the DOCwas kept constant in each testing mode.

The results show that the substrate heats up faster at a higherexhaust flow rate for similar fuelling modes. This can be explained

200

300

1000 1250 1500 1750 2000 2250Time [sec]

Te

Tg in

Mode 1 Mode 2 Mode 3 Mode 1

Fig. 38. Overall thermal response of aftertreatment (1400 RPM).

200

300

400

500

600

700

2650 3150 3650 4150

Time [sec]

Tem

pera

ture

[ o C

]

Engine speed: 1800 RPM Torque: 38 NmExhaust flow rate: 13.2 g/sExhaust O2: 11%

Tg in

DOC 1DPF 2

DPF 1

Mode 1 Mode 2 Mode 3 Mode 1

DOC 2

DOC DPF1 2 1 2Exhaust Flow

Fig. 40. Overall thermal response of aftertreatment (1800 RPM).

Table 3.9Energy analysis.

Engine speed (RPM), torque (Nm) Mode 1 Mode 2 Mode 3

1400, 37 1.716 1.518 1.5481800, 38 1.594 1.341 1.248

0

1

2

3

4

0 100 200 300Time [sec]

Rel

ativ

e fu

el in

ject

ion

rate

[%] Engine speed: 1800 RPM

Torque: 38 NmExhaust flow rate 13.2 g/sExternal fuel injectionParallel flow configuration

Mode 1

Mode 2

Mode 3

Mode 3: ΔT = 340ºC Mode 2: ΔT = 328ºC

Mode 1: ΔT = 351ºC

Fig. 39. Modes of external fuel injection (1800 RPM).


by the fact that convection heat transfer rate and oxygen supplyrate increases at higher exhaust flow rates.

3.2.1. Energy efficiency analysisAn energy efficiency analysis has been carried out for the

empirical results presented above. Due to the difference in thetemperature rise (DT) for the two exhaust flow rates under consid-eration, the specific energy consumption is defined by the authorsfor comparing different cases [15]. Fig. 41 shows the energy effi-ciency comparison between various modes of operation and theabsolute energy consumed to raise and maintain the temperaturerise of (DT) in the aftertreatment given in Table 3.9. It is observedthat the specific energy consumed is less at the high flow rate thanat the low flow rate. At 1800 RPM, due to the increase exhaust flowrate the rate of heat transfer increases, which may be attributed to

0

1

2

3

4

5

6

7

8

1 2 3Fuel Injection Modes

Spec

ific

ener

gy c

onsu

med

[kJ/

o C]

RPM 1400, Torque 37 NmRPM 1800, Torque 38 Nm

External fuel injection tests

Fig. 41. Comparison of different modes of external fuelling.

faster reaction rates. However, it can be seen that the substratecools down quickly at 1800 RPM compared to 1400 RPM. At higherflow rates, the convection heat transfer rate increases between theexhaust gas and the substrate. Thus, the thermal retention ability isreduced at higher flow rates.

At 1400 RPM, mode 2 gives the highest efficiency, whereas at1800 RPM, mode 3 results in the highest efficiency. Observationsreported in previous publications by authors also show a similartrend of energy consumption [16]. At a high fuelling rate, the ex-haust gas becomes rich in combustibles and therefore has lesschance to get oxidized completely. The time and temperature re-quired to the complete oxidation of the rest of combustibles insidethe DOC can be optimized by a combination of two or more fuelinjection rates during the heating process.

4. Concluding remarks

A one-dimensional DOC and DPF model has been developed inorder to perform simulations of active flow diesel aftertreatmentconfigurations. Simulations are carried out in order to investigatevarious active flow conditions for DPF regeneration. The simulationresults are compared with a set of experimental data for DPFregeneration. Based on the theoretical investigations, energy effi-ciency analysis is carried out between various active flow controlaftertreatment operations. A preliminary energy efficiency analysishas also been carried out with the empirical data of various activeflow aftertreatment control schemes.

Theoretical investigation indicated that soot oxidation insideDPF requires considerable less supplemental energy at the exhausttemperatures exceeding ‘‘light off” for DOC operation with externalfuel injection as compared to lower exhaust temperature. A sup-plemental exhaust gas heater can effectively reduce overall energypenalty to efficiently regenerate DPF. This may be attributed to thehigher thermal availability in the process of electrical heatingwhen compared with the exothermic heat release due to fueloxidation.

Simulation results indicate that at low-exhaust flow-rate oper-ation regimes, the DPF regeneration requires less supplemental en-ergy, providing that there is enough oxygen left in the exhaust forcompleting the soot oxidation reaction inside DPF channels. At theconditions of exhaust gas stagnation (e.g., no exhaust flow), thesoot oxidation may seize to occur due to the unavailability of oxy-gen inside the DPF.

Depending on the engine operating modes, the exhaust oxygenconcentration may vary in the exhaust stream. At low exhaust flowrates, variations in exhaust oxygen concentration may significantlyaffect the oxidation reaction efficiency of a DOC–DPF system.

Particulate regeneration is dependent on the factors such as theavailability of oxygen in the soot layer due to the exhaust wall flowand the exhaust temperature. It is observed from the simulationresults that soot profile gives an indication of the local soot oxida-tion rate. It is also observed that the higher the local soot oxidationrate, the higher is the tendency of substrate overheating and localthermal stress.

Variations across the channel length are neglected in modelconsideration. It is observed that with the results presented in thispaper, large portion of exhaust pressure drop across DPF is attrib-uted due to the soot layer distribution profile across DPF channel.


In active aftertreatment control strategies, critical supplementalenergy may be used to maintain the exhaust temperature abovethe ‘‘light-off” temperature, at those locations where the catalystmay exhibit best performance. Thus, the catalyst coating can bestrategically placed at most reactive regions thereby reducingoverall cost in precious metal coating.

From the empirical investigation it is observed that if engine ex-haust contains substantial amount of combustible substances,overheating can be predicted and prevented by increasing the flowrate with unidirectional flow control. The energy utilization effi-ciency of the external fuel injection can be optimized with boththe engine operating condition and substrate-operating conditionsincluding external fuelling strategies.

Acknowledgements

The authors are grateful for support from the University ofWindsor, the Canada Research Chair Programme, the Natural Sci-ences and Engineering Research Council of Canada, the CanadianFoundation for Innovation, Ontario Innovation Trust, AUTO 21TM(a member of the Network of Centres of Excellence of Canada pro-gramme) and the Ford Motor Company of Canada.

References

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