Comparison of CFD and LP Codes for the Simulation of ... file809.1 Comparison of CFD and LP Codes for the Simulation of Hydrogen Combustion Experiments Tadej Holler1, Ed M. J. Komen2,

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Comparison of CFD and LP Codes for the Simulation of Hydrogen Combustion Experiments

Tadej Holler1, Ed M. J. Komen

2, Ivo Kljenak

1

1 Jožef Stefan Institute (JSI)

Jamova cesta 39

SI-1000, Ljubljana, Slovenia

[email protected], [email protected]

2 Nuclear Research and Consultancy Group (NRG)

Westerduinweg 3

1755 ZG, Petten, The Netherlands

[email protected]

ABSTRACT

With computational advances, many tools are available for researchers and engineers to

allow different approaches for their investigative work in nuclear safety. When considering

the hydrogen threat in severe accident analyses in nuclear power plants, particularly hydrogen

deflagration, mainly two approaches are being used, i.e. Lumped Parameter (LP) and

Computational Fluid Dynamics (CFD) codes. In this paper, a comparison of those two

approaches is performed based on three experiments performed in the THAI experimental

facility.

The results shown in this work indicate that more detailed analysis is obtainable using

the CFD approach. However, the LP approach is much better in obtaining some general

information about the hydrogen deflagration events while at the same time requiring

substantially less computational resources.

1 INTRODUCTION

During a postulated severe accident in a Light Water Reactor (LWR) Nuclear Power

Plant (NPP), hydrogen can be formed from steam, e.g. from zirconium-steam interaction [1].

When it is mixed with the surrounding air in the containment, it can form a highly flammable

mixture. Hydrogen combustion in NPP containment therefore poses a direct threat to the

relevant safety systems, structures and components and, most importantly, can even

compromise the containment integrity itself [1]. Such scenarios of hydrogen generation and

combustion inside a LWR NPP containment were confirmed in the Three Mile Island NPP

accident in the USA in 1979 [2] and in the Fukushima Daiichi NPP accident in Japan in

2011 [3]. In the latter, the destructive power of hydrogen combustion was demonstrated.

Despite the hydrogen mitigation systems installed to reduce the hydrogen risk, such as

Passive Auto-catalytic Recombiners (PARs) and igniters, reliable hydrogen risk assessment is

still required.

Mainly two types of codes are used for hydrogen safety analyses: a) lumped parameter

(LP) codes, e.g. ASTEC [4], CONTAIN [5], MELCOR [6], and SPECTRA [7] and b)

Computational Fluid Dynamics (CFD) codes, e.g. ANSYS Fluent [8], ANSYS CFX [9] and

COM3D [10]. LP codes provide relatively quick estimates of the pressure loads, while CFD

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Proceedings of the International Conference Nuclear Energy for New Europe, Portorož, Slovenia, September 5-8, 2016

codes are much more CPU demanding, but offer a local-scale representation of the

phenomena involved.

In the nuclear research community, there is currently an ongoing debate about whether

CFD is really necessary to predict hydrogen deflagration phenomena occurring in real-scale

containments. This issue is often raised, since Lumped Parameter (LP) codes can predict the

mean maximum pressure 𝑝𝑚𝑎𝑥 or the Adiabatic Isochoric Complete Combustion (AICC)

pressure 𝑝𝐴𝐼𝐶𝐶 well in case of hydrogen combustion. However, during fast deflagrations,

intermediate peak pressures occur, which can be substantially higher than 𝑝𝐴𝐼𝐶𝐶. Estimation

of these maximum pressures is required in order to assess possible structural damage due to

hydrogen combustion.

Furthermore, the evaluation of possible structural damage of the containment does not

involve solely the maximum pressure, but the dynamic loads, e.g. rate of pressure change

versus time, must be considered as well. That is, in addition to the maximum pressure 𝑝𝑚𝑎𝑥 in

the containment, the rate at which the pressure increases with time, i.e. 𝑑𝑝 𝑑𝑡⁄ is important.

Also, the containment can undergo structural damage, if the residual pressure that remains

after complete burn oscillates with eigen frequencies corresponding to those of the

containment structure or the safety systems inside the containment. Therefore, both the

frequency and the amplitude of the pressure oscillations may also be important. This

essentially means that four parameters are important for the evaluation of the possible

structural damage: a) maximum pressure, b) rate of pressure increase 𝑑𝑝 𝑑𝑡⁄ , c) eigen

frequencies, and d) amplitudes corresponding to the residual pressure waves. The latter three

parameters cannot be obtained using an LP code, since these parameters depend on the flame

acceleration. This flame acceleration in turn depends on the generation of turbulence in the

flame front and the interaction of this flame with obstacles.

To show the importance of the effect of obstacles present, let us consider for example

experimental data obtained from the THAI and ENACCEF facilities. Such data were released

for CFD benchmarking in the OECD International Standard Problem No. 49 [11]. For a

similar gas composition, e.g. around 13 vol. % hydrogen concentration, the maximum flame

propagation velocity achieved in the THAI facility [12] was about 10 m/s, while in the

ENACCEF facility the maximum flame velocity reached almost 150 m/s (see Figure 1). That

is about an order of magnitude difference in the flame speed for almost the same gas

composition. This large difference in the flame propagation velocity is due to the turbulence

generated by the obstacles present in the ENACCEF facility. This turbulence generation leads

to increased flame acceleration.

Such flame acceleration due to turbulence generation by local obstacles cannot be

computed using LP codes. Furthermore, the propagation of the pressure wave phenomena

induced by the flame acceleration also cannot be computed by LP codes. Moreover, this

example demonstrated that no conclusions concerning flame propagation can be made based

on the evaluation of the gas mixture composition only. Basically, this is done when

conclusions are based solely on the usage of the frequently used Shapiro diagram [13]. The

considered local effects cannot be computed by LP codes, but can be computed by CFD

codes. However, such detailed CFD analyses consume substantially more resources than the

LP approach.

This paper focuses on comparison of results obtained only on experiments performed in

slow deflagration regime and described in more detail in the next section.

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Figure 1: Axial flame propagation velocity observed in

THAI and ENACCEF experimental facilities.

2 EXPERIMENT

The THAI containment test facility, shown in Figure 2, is operated by Becker

Technologies GmbH in Eschborn, Germany. It is a vertical cylindrical stainless steel vessel

designed to withstand pressures of 20 bar at wall temperature 20 °C. The facility is measuring

9.2 m in height and 3.2 m in diameter, bringing the total volume to 60 m3. Within the

framework of the hydrogen deflagration (HD) part of the OECD-NEA THAI project [11,12],

29 different experiments were performed, aiming to provide a deeper insight into the

phenomenology of hydrogen deflagrations. In these experiments, varying initial conditions

were used, by changing the initial hydrogen and steam concentrations, temperature, pressure,

ignition location, and atmosphere stratification. During these experiments, all inner structures

that could possibly induce turbulence were removed, except for the experimental

instrumentation, which consisted of 43 thermocouples and 4 pressure transducers distributed

throughout the vessel.

For the purpose of this paper, three experiments with bottom ignition and different

hydrogen concentrations were selected. The initial conditions of these experiments are

presented in Table I.

Table 1: Initial conditions for considered THAI HD experiments

Hydrogen

concentration

[vol.%]

Temperature

[K]

Pressure

[kPa]

Mixture

THAI HD-12 8.0 291.0 148.5 Uniform

THAI HD-3 9.0 295.5 148.5 Uniform

THAI HD-7 9.9 290.0 148.0 Uniform

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3 PHYSICAL AND NUMERICAL MODELING

3.1 Lumped Parameter Code – ASTEC

3.1.1 Physical modelling

The ASTEC code [6] is being jointly developed by the Institut de Radioprotection et de

Sûreté Nucléaire (IRSN – France) and the Gesellschaft für Anlagen- und Reaktorsicherheit

(GRS – Germany). To simulate hydrogen deflagration, the FRONT model in the ASTEC code

calculates the flame propagation from one control volume (into which the entire vessel is

divided during nodalisation) into adjacent ones. The flame propagation takes place within the

junctions of the system (flow connections between the volumes). The combustion of the gas

mixture takes place in the volumes and is calculated by a combustion model. The flame

propagates through junctions. At the end of each junction, it is checked whether the

conditions in the target zone are prone to ignition, based on the Shapiro diagram. In case the

ignition is possible, it occurs instantaneously or by user controlled boundary conditions – by

time or hydrogen concentration. After the combustion completeness criteria are reached, the

combustion stops. The laminar burning velocity is calculated according to the Liu-

MacFarlane correlation [14]. Turbulent flame front velocity is calculated via the Peters

correlation [15].

3.1.2 Nodalisation

In the input model for the ASTEC code, the THAI vessel was divided into 16 control

volumes (Figure 3). This nodalisation is a compromise between a much lower number, which

would not enable to simulate the gradual flame propagation in the axial direction, and a much

higher number of small volumes, for which lumped parameter codes were not developed.

Figure 2: Schematic of the THAI

experimental vessel.

Figure 3: Nodalisation of THAI

vessel for ASTEC code.

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3.2 CFD Code – ANSYS Fluent

3.2.1 Physical modelling

The ANSYS Fluent CFD code [8] was used as a platform on which the combustion

model was implemented. The density-based coupled solver was applied together with an

explicit time stepping method. All spatial discretization schemes used were of second order.

The applied turbulence models were the standard 𝑘 − 𝜀 models.

In addition to the equations for conservation of mass, momentum and energy, the

additional transport equation for the Favre-averaged progress variable, �̃�, is resolved. This

progress variable is defined as:

�̃� =�̃�𝑓 − 𝑌𝑓,𝑢

𝑌𝑓,𝑏 − 𝑌𝑓,𝑢 (1)

Here, 𝑌𝑓 is the mass fraction of fuel, with indexes 𝑢 and 𝑏 representing the unburned

and burned states, respectively. Lipatnikov and Chomiak [16] proposed the following

progress variable transport equation for the Flame Speed Closure (FSC) combustion model:

𝜕�̅��̃�

𝜕𝑡+

𝜕�̅��̃��̃�𝑗

𝜕𝑥𝑗=

𝜕

𝜕𝑥𝑗[�̅�(𝜅 + 𝐷𝑡)

𝜕�̃�

𝜕𝑥𝑗] +

(𝑆𝑙)2

4(𝜅 + 𝐷𝑡)𝜌𝑢�̃�(1 − �̃�) + 𝜌𝑢𝑈𝑡|∇�̃�| (2)

In equation (2), 𝜅 and 𝐷𝑡 are molecular and turbulent diffusivities, respectively, 𝜌𝑢 is

the unburned gas density and 𝑆𝑙 represents the laminar flame speed. The third term on the

right hand side of the equation (2) is the quasi-laminar source term for progress variable, with

which the Turbulent Flame Speed Closure (TFC) combustion model [17], developed by

Zimont for fully turbulent premixed flames, was expanded by Lipatnikov and Chomiak. The

turbulent burning velocity 𝑈𝑡 is modeled, based on the work by Zimont, as follows:

𝑈𝑡 = 𝐴𝑢′𝐷𝑎1 4⁄ 𝐹(𝐿𝑒) (3)

Here, 𝐴 is a model constant, 𝑢′ is the turbulent intensity and 𝐷𝑎 is the Damkohler

number. It should be noted here that the model constant 𝐴 = 0.4 was the same for all the

considered computations. This follows the procedures from the previous FSC model

validations [18-20]. The preferential diffusion thermal (PDT) instabilities were accounted for

using the 𝐹(𝐿𝑒) term [18]. In the planar, fully developed case, the added source term results

in the simple expression for the flame speed, 𝑆𝑡:

𝑆𝑡 = 𝑆𝑙 + 𝑈𝑡 (4)

3.2.2 Mesh and boundary conditions

To obtain accurate results, while maintaining computational times at reasonable levels,

an adaptive mesh refinement (AMR) technique was applied to the flame front. A grid

sensitivity study was performed in an earlier work [18-20], which indicated that using two

levels of AMR produces practically grid independent results. Each level of refinement means

halving the size of a computational cell in every direction of the axisymmetric 2D model used

in our calculations. The base mesh cell size was around 2 cm by 2 cm, while the ignition

region was further refined to contain cells measured approximately 0.6 mm by 0.6 mm, to

keep the shape of the ignition region as even as possible. The ignition, located 1.55 m above

the floor of the vessel, was modelled with an ignition sphere with 8 cm radius, to which

burned properties were prescribed. The flame front position was extracted from the

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coordinates corresponding to the iso-surface value of c = 0.5, as done in our previous work

[18-20]. In the experiment, the reactants in the experiment were mixed using recirculation

fans and then switched off several minutes prior to the ignition. Thus, the initial turbulence

values were assumed to be low, although no measurements of these parameters were

performed. Initial values for k and ε were selected as 10-4 m2s-2 and 4.8·10-5 m2s-3,

respectively, throughout the computational domain. This is also done in accordance with the

previously conducted validations with the CFD-based modelling [18-20].

For slow deflagration computations in our previous work [18-20] it was shown that

complete heat transfer mechanism, including radiation, has a significant impact on total

energy levels and consequently on the pressure build-up. Therefore, the simulation results

were computed using prescribed constant wall temperature along with the Discrete Ordinates

(DO) thermal radiation model [8].

4 RESULTS AND DISCUSSION

With increasing initial hydrogen concentration, the trends of increasing pressure peak,

faster flame propagation and consequently higher pressure increase rate were observed in

experiments. Obtained simulation results with both LP and CFD approach are compared to

corresponding experiments HD-12, HD-3, and HD-12, by observing the pressure, axial flame

propagation and pressure increase rate 𝑑𝑝 𝑑𝑡⁄ .

4.1 Pressure

From the point of view of nuclear safety, the accurate prediction of the maximum (peak)

pressure that is reached during combustion is essential. Experimental and calculated pressures

are shown in Figures 4, 5, and 6, for cases HD-12, HD-3 and HD-7, respectively.

The overall observation is that both LP and CFD simulations provided adequate results

of maximum pressure. The trend of maximum pressure increase with increasing initial

hydrogen concentrations, as observed in the experimental results, was also reproduced

qualitatively well by both LP and CFD codes. Relatively mild pressure oscillations were

observed in the experiments with higher initial hydrogen concentrations, i.e. HD-3 and HD-7,

which were not captured by the CFD approach based on the described combustion modeling,

due to numerical dissipation (and were not expected to be captured by LP codes anyway).

However, it should be noted here, that the CFD modelling approach has already shown to be

capable of capturing the pressure oscillations during combustion [20].

Figure 4: Experimental and calculated pressure for run HD-12 (8.0 vol.% H2).

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4.2 Axial Flame Propagation

Experimental and calculated axial flame propagations are shown in Figures 7, 8, and 9,

for cases HD-12, HD-3 and HD-7, respectively. A trend of flame propagation rate increase

with increasing initial hydrogen content can be observed in the experiment. However, some

slowdown of flame propagation can be observed in CFD simulations with the flame

approaching the top of the vessel, especially in Figure 9. This rather peculiar flame behavior

can be attributed to the movement of the flame front towards the side of the vessel. The

observed “step” indicates the moment the flame touches the side wall, after which the

movement of the flame is again more intense in the axial direction.

As to the results obtained with the LP code, they are similar for all three considered runs

and, what is more important, very different from the experimental curves. This is a

consequence of the very simple modelling of flame propagation using the LP approach.

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Figure 7: Experimental and calculated axial flame propagation

for run HD-12 (8.0 vol.% H2).





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4.3 Pressure Increase Rate

For an easier interpretation of the pressure increase from pressure history results,

experimental and calculated pressure increase rates are shown in Figures 10, 11, and 12, for

cases HD-12, HD-3 and HD-7, respectively. Namely, both the maximum pressure and the

pressure increase 𝑑𝑝 𝑑𝑡⁄ development are considered highly important parameters for analysis

of the structural integrity of the NPP containment.

For the case HD-12, the CFD simulation overpredicts the maximum 𝑑𝑝 𝑑𝑡⁄ compared to

the experimental results for almost 50%, whereas the LP simulation overpredicts it only

slightly. However, in cases with higher initial hydrogen concentrations, i.e. HD-3 and HD-7,

the CFD simulations underpredict the experimental results by around 50% and 30%,

respectively. In the case HD-3, the LP simulation also underpredicts the maximum (albeit less

than the CFD simulation), whereas in the case HD-7, it overpredicts it.

As already shown in Figures 4-6, the simulations are capable of capturing the trend of

the earliest occurrence of first pressure increase, with the increase of initial hydrogen

concentration. For CFD simulations, this cannot be said for the trend of maximum 𝑑𝑝 𝑑𝑡⁄

values, since this parameter in simulation run of HD-12 is reached before the maximum of

case HD-3, which has a higher initial amount of hydrogen. For LP simulations, these

differences are too small to be considered.

Figure 10: Experimental and calculated pressure increase rate 𝑑𝑝 𝑑𝑡⁄


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5 CONCLUSIONS

For the case of slow hydrogen deflagration, considered in the present paper, results of

pressure and pressure increase rate calculated with the LP code are just as adequate, if not

even more, than results calculated with the CFD code. From the point of nuclear safety,

simulations with LP code appear to be sufficient, making lengthy simulations with CFD codes

unnecessary.

However, the comparison of calculated flame propagations with LP and CFD codes

clearly confirms, that LP codes do not simulate the basic mechanisms of combustion. Thus,

the reliability of results calculated with LP codes may be questionable for conditions for

which they have not been validated due to lack of suitable experimental data. In such cases,

CFD codes could be trusted more, as they solve the basic equations of fluid mechanics. Thus,

in the long term, the authors recommend the further development of CFD codes to improve

their predictions of pressure increase rate as well as to increase their calculation speed by

using efficient numerical methods.

ACKNOWLEDGMENTS

The work described in this paper was funded by the Dutch Ministry of Economic

Affairs, the Slovenian Research Agency (Research programme “Reactor Engineering”) and

the European Commission (CESAM project, 2013-2017).

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Documents

Comparison of CFD and LP Codes for the Simulation of ... file809.1 Comparison of CFD and LP Codes for the Simulation of Hydrogen Combustion Experiments Tadej Holler1, Ed M. J. Komen2,