Elise El Ahmar

Jean-Yves Coxam

Christophe Coquelet

Karine Ballerat-Busserolles

Excess molar volumes and enthalpies of binary mixtures containing furan and toluene, 1-octanol or nhexane at temperatures up to 383.15 K and pressure up to 1.6 MPa

Keywords: furan, toluene, 1-octanol, n-hexane, experimental excess properties, modelling

Densities of pure furan, toluene, 1-octanol, n-hexane and, of binary liquid mixtures containing furan, has been measured from 303.15 K to 383.15 K at pressures up to 1.6 MPa, using a vibrating tube densimeter. Excess molar volumes have been calculated from the densities with an average standard uncertainty 0.02 cm 3 .mol -1 . Excess molar enthalpies of the binary mixtures were measured at temperatures ranging from 313.15 to 383.15 K at 0.5 MPa and/or 1.0 MPa, using flow mixing calorimeter. The average standard uncertainty on the whole measurements is less than 1%. The excess molar properties derived from experimental data were correlated using a Redlich-Kister polynomial functions, NRTL activity coefficient model and Prigogine Flory Patterson modelling is described and used to correlate the data.

Introduction

Due to the depletion of fossil resources, green chemistry has now become a real alternative to petrochemicals.

Lignocellulosic biomass (LCB) refers to plant biomass that is composed mainly of lignin and cellulose (wood residues, cereal straw and forestry waste) [START_REF] Harmsen | Literature Review of Physical and Chemical Pretreatment Processes for Lignocellulosic Biomass[END_REF]. Currently, this biomass is the most important source of renewable energy and the only renewable source of carbon. It can provide for about 10% of total energy consumption worldwide (IEA Statistics, 2021). However, this cannot be considered as the optimal use of LCB resources because much of this consumption concerns firewood-based heating and cooking. This biomass is also playing a major role in the chemical sectors as a substitution material. Nevertheless, the major part of the annual production of biomass (220 billion tonnes per year) is still underexploited and, there remains considerable scope for the development of LCB value chains for the production of energy, chemicals, polymers and materials.

Molecules derived from lignocellulosic biomass pyrolysis belong to a variety of complex oxygen compound families [START_REF] Kawamoto | Strong interactions during lignin pyrolysis in wood -A study by in situ probing of the radical chain reactions using model dimers[END_REF]. The main families identified are carbonyls, sugars, oxygenated aromatics and furans. The thermodynamic behavior of these molecules is not well known because the models and methods traditionally used for hydrocarbons fail to return it. Indeed, these molecules are the site of complex molecular interactions. Consequently, experimental determination of excess thermodynamic properties will be necessary. These properties will be required during new product design chemicals and development of new production processes from lignocellulosic material. This paper focus on the furan as a derived molecule obtained from the pyrolysis of the LCB. Furan can be used as a precursor for the synthesis of pharmaceutical molecules [START_REF] Joule | Heterocyclic Chemistry[END_REF] as well as fongicides, herbicides or disinfectants [START_REF] Fredus | Industrial Uses of Ind[END_REF]. To the best of our knowledge, only few data for Vapour-Liquid and Liquid-Liquid Equilibrium of binary and ternary mixtures containing furan are available in the open literature ( [START_REF] Nala | Vapour-liquid equilibrium (VLE) for the systems furan+n-hexane and furan + toluene, measurements,data treatment and modeling using moleculat models[END_REF], [START_REF] Auger | Equilibrium data and GC-PC SAFT predictions for furanic extraction[END_REF] and [START_REF] Théveneau | Vapor liquid equilibrium data for the Furan -Toluene binary system between 313[END_REF]). Apparent Henry's law constants of Furan in various n-Alkanes and alcohols were obtained by Hajiw et al. [START_REF] Hajiw | Apparent Henry's law constants of furan in different n-alkanes and alcohols at temperatures from 293 to 323 K[END_REF] and, the excess enthalpy for the binary mixture {n-hexane-furan} were measured at 298.15K and 303.15K by M. D. Guillen [START_REF] Guillen | Excess enthalpies and excess volumes of n-hexane and of tetrachloromethane with furan, 1,4-dioxane, tetrahydrofuran, and tetrahydropyran[END_REF]. As a result, a study of excess properties has been realized for the binary mixtures constituted with furan and n-hexane, toluene or 1-octanol.

These molecules were selected as potential solvent for liquid-liquid extraction.

Experimental

Chemicals.

The chemicals used are listed in table 1 together with molecular formula, CAS number, suppliers, purity given by the supplier and molecular weights. All the compounds were used without further purification.

Properties studied

In this paper, two excess molar properties have been investigated: excess molar volumes and excess molar enthalpies. The excess property Y E , Eq. ( 1), represents the difference between the value Y measured in the real solution and the value Y id calculated for an ideal solution, at the same conditions of pressure, temperature, and composition, [START_REF] Privat | Discussion around the paradigm of ideal mixtures with emphasis on the definition of the property changes on mixing[END_REF]. To measure excess properties, the system will have to be in liquid phase before and after mixing.

The normal boiling points are: 304.2 K for furan, 383.5 K for toluene and, 341.8 K for n-hexane [11]. To remain in liquid phase at investigated temperatures [303. 15 K -383. 15 K], our experiments were performed at 1.65 MPa for the densimetric measurements and, at 0.5 and 1.0 MPa for the enthalpy measurements.

Y E = Y-Y id (1)
Preparation of the binary mixtures.

For densimetric measurements, 45 g to 50 g of binary mixtures were prepared by weight in Schott Duran flasks (50 ml) using a Mettler Toledo AE163 balance with a precision of 0.1 mg. the remaining gas space after preparation was flushed with nitrogen before hermetically closing the flasks. The uncertainty on the molar fraction is estimated to be better than 10 -4 .

For flow calorimetric measurements, binary mixtures are prepared directly in situ. The uncertainty on the composition is obtained from a calculation of propagation of error including the uncertainty on the density and on the volume flow rate of pure components. The relative uncertainty on volume flow rate, using an ISCO pump (Model 100 DM) is 0.03%. Considering an uncertainty of 10 -2 kg.m -3 on pure component density, the final uncertainty on the molar fraction of the mixture is estimated to be 2.10 -4 .

Volumetric technique

Densities were obtained at 3 temperatures, namely 303. 15 The density of mixtures, m, is derived from periods of vibration using Eq. ( 2)

) ( 2 2 ref m ref m K         (2)
where m and ref are the period of vibration when flowing solution (m) or reference fluid (ref) respectively. ref  is the density of the reference fluid and K is a calibration constant. This calibration constant K was determined by measurements of water and NaCl solutions with well-known densities. The relative uncertainty in the calibration constant is about 0.3%. The reference density of water was calculated with the equation of Hill [START_REF] Hill | A unified fundamental equation for the thermodynamic properties of H2O[END_REF], the reference density of NaCl solutions was obtained from the recommendations of Archer [START_REF] Archer | Thermodynamic properties of the NaCl-H2O system. II. Thermodynamic properties of NaCl(aq), NaCl •2H2O(cr), and phase equilibria[END_REF]. The stability of the measured periodsm and ref was typically 5.10 -6 ms at 303.15 K and 1.10 -5 ms at higher temperatures. After calculation of propagation of error, the uncertainty in the density is found to be about 10 -2 kg.m 3 .

The density of pure furan was measured with water as reference fluid considering Eq. ( 2). For the other pure compounds as well as for mixtures, the reference fluid was furan. The measurements are performed at a constant flow rate of 0.1ml.min -1 . The overall setup of the experiment is represented in the Fig. 1. Furan is first circulating into the vibrating tube densimeter using an DM 100 ISCO high pressure syringe pump (flow line (1)). The second pure compound (or the mixture) is introduced in a sample loop prior to experiment. After registering the value of the period of vibration for furan (ref), the second fluid (or mixture) is then introduced (flow line (2)), and the period of vibration (m) is registered. Excess molar volumes V E are calculated from the experimental density data using Eq. (3)

2 2 2 2 2    M x M x M x M x V fur fur fur fur fur E           (3)
where x is the molar fraction,  the density and M the molar mass; subscript "fur" refers to furan and "2" refers to the second pure compound. The estimated average total uncertainty, including molality calculations, and vibrating tube period noise and reproducibility on the excess molar volume is 0.01 cm 3 .mol -1 .

Calorimetric technique

The excess molar enthalpies are obtained by measuring thermal effect due to the mixing of furan with a second compound, using a Setaram BT2.15 heat conduction differential calorimeter. The fluids mix in a flow mixing cell developed originally in the laboratory [START_REF] Arcis | Enthalpy of solution of CO2 in aqueous solutions of 2-amino-2-methyl-1propanol[END_REF]. The thermal effect (heat power) is detected by a calibrated thermopile.

Measurements were carried out in flow mode at constant temperature and pressure. The excess molar enthalpies are measured for different molar fractions of furan, obtained by changing the molar flow rates ration ( 2 / n n fur  

). The two fluids (furan and compound 2) are supplied by two high-pressure syringe pumps (Isco Model 100 DM). These two pumps are regulated at a constant temperature of 303.15 K, using a JULABO thermostatic bath (model FP 89) that permit a temperature control with a stability of 0.1K. Molar fraction of furan is calculated from molar flow rates (𝑛̇) using Eq. [START_REF] Fredus | Industrial Uses of Ind[END_REF].

𝑥 𝑓𝑢𝑟 = 𝑛̇f ur 𝑛̇𝑓 𝑢𝑟 +𝑛̇2 with 𝑛̇= 𝜌•𝑣Ṁ (4)
The molar flow rates (𝑛̇) were calculated from the pump volume flow rates (𝑣) using the densities () and the molar mass of the fluids (M). The total volume flow rate arriving at the mixing point was maintain constant to a value of 0.5 ml.min -1 .

The molar excess enthalpy H E is derived from the heat power of mixing and the molar flow rates n  (mol.s -1 ) as indicated in Eq. ( 5); the calorimetric signal S (mV) is converted into heat power using a calibration constant K (mV•mW -1 ).

𝐻 𝐸 /𝐽𝑚𝑜𝑙 -1 = 𝛥𝑆 𝐾(𝑛̇𝑓 𝑢𝑟 +𝑛̇2) [START_REF] Nala | Vapour-liquid equilibrium (VLE) for the systems furan+n-hexane and furan + toluene, measurements,data treatment and modeling using moleculat models[END_REF] The calorimetric signal S represents the difference between the signal recorded during mixing and a baseline signal recorded when only furan is flowing into the circuit. This signal is varying between 100mV and 800 mV, depending on the binary mixture composition for furan + 1-Octanol and furan + n-hexane. The calibration constant K is determined using the reference data for excess enthalpies of {ethanol-water} system recommended by Ott et al. [15,[START_REF] Ott | Excess enthalpies for (ethanol + water) at 398[END_REF]. Considering the uncertainties on flow rates, thermopile calibration K and calorimetric signal stability and temperature fluctuations, the relative expanded uncertainty on enthalpies are lower than 5 % for very dissymmetric mixtures (large differences in compressibility or molecular weight of the compounds, or large differences in the flow rates of the pumps). For the compounds studied in this paper this uncertainty is reduced to 1%.

Results and discussions

In this section, the density, the excess molar volume and the excess molar enthalpy of binary mixtures will be presented at different pressure and temperature. The excess molar properties (volume and enthalpy) will be correlated using the Redlich-Kister (RK) polynomial function and the NRTL activity coefficient model as implemented in the Simulis thermodynamic software developed by Prosim® (France).

Density of pure compounds

The density and molar volume of pure compounds at 303.15 K and pressures up to 2.1MPa are presented in Table 2.

Table 2 Densities ( and molar volumes (V M ) of pure compounds at 303.15 K and pressures up to 2.1MPa Standard uncertainties u are: u()= 2.10 -2 kg.m 3 for furan, n-hexane, toluene and 1-octanol at 303.15 K; u()=5.10 -2 kg.m 3 for 1-octanol at 353 K; u()= 7.10 -2 kg.m 3 for 1-octanol at 383 K; u(V M ) = 0.01 cm 3 •mol -1

T / K p /MPa  / kg.m -3 V M / cm 3 .
The molar volumes were fitted to a Tait's equation [START_REF] Dymond | The Tait equation: 100 years on[END_REF] as presented in Eq. ( 6); the parameters v(0,T), B(T) and C calculated at 303. 15 K are reported in Table 3.

𝑉 𝑀 (𝑝, 𝑇) = 𝑣(0, 𝑇) {1 -𝐶 𝑙𝑛 [1 + 𝑝 𝐵(𝑇) ]} (6)
Table 3 Parameters for Tait's equation [START_REF] Dymond | The Tait equation: 100 years on[END_REF], at 303.15 K.

Compound

(0, ) vT / cm



The densities at 0.1 MPa reported in Table 2 were derived from Eq. ( 6) to be compared to available literature values.

More than 30 references are available for the three compounds studied. The minimum and the maximum literature values for density are reported in the Table 2, at 303.15 K and 0.1 MPa. In all cases our experimental value is included in the maximum-minimum interval. The molar volume of pure compounds increases mostly linearly in our temperature domain; the evolution is more important for n-hexane.

Experimental results for the binary mixtures

Furan + n-hexane

Because of furan volatility, it was necessary to run experiment above 0.1 MPa. In order to test the influence of pressure on excess enthalpies, measurements at 353. 15 4 and represented in Fig. 2. As it was shown that there is no influence of the pressure on excess enthalpies within the range studied, values obtained at 0.5 MPa and 1.0 MPa were considered as a single set of data for RK correlation at 353.15 K.

Table 4

Volumetric and calorimetric data obtained for the binary mixture furan + n-hexane. Densities () and excess molar volumes (V E ) of binary mixture furan + n-hexane at 1.6 MPa. 3 .mol -1 ) 303. 15 A Redlisch-Kister (RK) correlation was first use to describe the excess molar properties for each temperature using Eq. [START_REF] Théveneau | Vapor liquid equilibrium data for the Furan -Toluene binary system between 313[END_REF].

x furan  (kg.m -3 ) V E (cm 3 .mol -1 ) x furan  (kg.m -3 ) V E (cm 3 .mol -1 ) x furan  (kg.m -3 ) V E (cm
𝑋 𝐸 = 𝑥 2 (1 -𝑥 2 ) ∑ 𝐴 𝑛 (1 -2𝑥 2 ) 𝑛 (7)
Where X is Volume or Enthalpy, An are the adjustable coefficients of the equation, and n the number of parameters to be considered. The parameters are given in table 5 and Table 6.

The average standard error of the estimates is 0.02 cm 3 .mol -1 for the volumetric properties and less than 3% for the calorimetric properties. From these experimental data and those found in the open literature [START_REF] Guillen | Excess enthalpies and excess volumes of n-hexane and of tetrachloromethane with furan, 1,4-dioxane, tetrahydrofuran, and tetrahydropyran[END_REF], it has been noticed that for this polar -nonpolar mixture, the excess molar enthalpies and volumes are positive for the entire composition range and for all temperature studied (Fig. 2). In fact, the attractive interactions between molecules of different compounds are, on average, weaker than those between molecules of the same compound. The excess molar enthalpy is a decreasing function of the temperature; meanwhile, the excess molar volume is an increasing function of the temperature.

Furan + 1-octanol

The densities, molar volumes and enthalpies of the binary mixture furan + 1-octanol are presented in Table 7, and plotted in Fig. 3. The experimental excess enthalpies are positive for the entire composition range. Table 7 Volumetric and calorimetric data obtained for the binary mixture furan + 1-octanol.

Densities () and excess molar volumes (V E ) of binary mixture furan + 1-octanol at 1.6 MPa.

x furan  (kg.m -3 ) V E (cm 3 .mol -1 ) x furan  (kg.m -3 ) V E (cm 3 .mol -1 )
x furan  (kg.m -3 ) V E (cm 3 .mol -1 ) 303. 15 Meanwhile, for the excess molar volume, we notice that even so the values are weak, and lower than 0.4 cm 3 .mol -1 , they change from positive to negative when increasing temperature. This can be explained by a change in the molecular interactions in the solution. At low temperature, the interactions between molecules of same species are stronger while at the highest temperature, intermolecular interactions are predominant, leading to a contraction of the solution. This is probably linked with the change in the evolution of the excess enthalpy that goes through a maximum before decreasing at higher temperature. The RK correlations were also realized for this system, with 3 parameters for both properties. However, for the excess molar volumes at 353.15 K, due to the wave form of the curve, the standard state deviation is poor and a fit with 4 parameters was needed.

Furan + toluene

The excess molar volume and enthalpies of the binary mixture furan + toluene are presented in Table 8 and plotted in Fig. 4.

Table 8 Volumetric and calorimetric data obtained for the binary mixture furan + toluene.

Densities () and excess molar volumes (V E ) of binary mixture furan + toluene at 1.6 MPa.

x furan  (kg.m -3 ) V E (cm 3 .mol -1 ) x furan  (kg.m -3 ) V E (cm 3 .mol -1 )
x furan  (kg.m -3 ) V E (cm 3 .mol -1 ) 303. 15 Standard uncertainties u are u(V E ) =0.01 cm 3 .mol -1 ; u(x) =0.0001.

For furan + toluene, the experimental excess enthalpies are positive for the entire composition range and their values are very weak with a maximum value 50 J.mol -1 . The experimental excess molar volumes are close to zero (lower than 0.1 cm 3 .mol -1 ). This behaviour indicates that the properties of the mixture are close to the ideal properties.

The RK correlation was done with 3 parameters, but they are not significant due to the weakness of the excess properties. The lines are RK correlations with 3 parameters.

NRTL correlation of excess molar enthalpies

The NRTL activity coefficient model was applied for binary mixtures furan + n-hexane and furan + 1-octanol. The binary mixture furan + toluene was not considered as the excess enthalpies are nearly equal to zero. The nonrandomness parameter  was adjusted to a constant value of0.17. The interaction parameters are temperature dependant and have been adjusted for each temperature using Simulis Thermodynamics software, from Prosim ® . As the pressure has no significant effect on the excess molar enthalpy and, as the parameters are not pressure dependant, the data obtained at 0.5 and 1.0 MPa for the same temperature were treated as a single set of data. The NRTL parameters are given in Table 9 for both systems furan + n-hexane and furan + 1-octanol.

Table 9 Interaction parameters of the NRTL activity coefficient model for excess molar enthalpies of binary mixtures furan + n-hexane and furan + 1-octanol with  = 0.17. A good agreement (Fig 5) between models and experimental results is observed. It can be seen that NRTL model is able to accurately represent the excess molar enthalpies of the two binary systems studied. The AAD is less than 1%

for both binary mixtures. The NRTL correlation will be used in the next part to compare the different binary mixtures at a given temperature.

Comparison of excess properties of binary mixtures at given temperature

The excess molar volumes and enthalpies are plotted in Fig. 5 as function of the furan molar fraction, at 303.15 K and 313.15 K for V E and H E , respectively. The correlation shows a good agreement with the data as mentioned in part 3.3. The thermodynamic excess properties are very close to zero for the binary mixture furan + toluene (Fig. 5).

The excess molar volume V E is larger for n-hexane + furan than for the other mixtures while the largest excess molar enthalpy is reached by furan + 1-octanol mixture. The maximum of the thermodynamic properties is displaced to the highest molar fraction of furan when it is mixed with 1-octanol: V E is maximum when xfuran~0.6 and, H E is maximum when xfuran ~0.7. This behaviour is linked to specific intermolecular interactions between the alcohol group and le furan, that tends to skewed the properties toward mixtures rich in furan. For the other systems the maximum is located around xfuran~0.5 because of more simple molecular interactions.

Prigogine Flory Patterson modelling

In order to analyse the excess molar enthalpy, it is possible to use the Prigogine Flory Patterson model [START_REF] Patterson | Corresponding states theories and liquid models[END_REF]. As the pressure effect is not significant in our experimental data, we apply the PFP model for the two binary systems furan + n-hexane and furan + 1-octanol. We do not use this model for the binary system furan + toluene due to the low value of the excess enthalpy. The particularity of this model is to decompose the excess property into several contributions. Concerning excess molar enthalpy, it exists mainly two contributions. We suppose that there is no chemical reaction between the chemicals. It is important to remind that this model was originally used to analyse the excess thermodynamic properties of non-polar mixture and was extended to polar mixtures. We can cite the recent work of Coquelet et al. [START_REF] Coquelet | Density and excess volume for four systems involving eugenol and furan[END_REF] on the excess volume of similar binary systems.

The PFP model describes the interaction between the chemical species and the expression of the excess molar enthalpy is reported in Eq. ( 8).

For the excess molar enthalpy, the two main contributions are:

-Interactional contribution: this contribution is proportional to the Flory parameter. This parameter can be calculated using solubility parameter and reflects the energy change during the contact between the unlike molecules. In previous work [20] it has been considered that this parameter can be expressed as a function of the volume fraction.

-Free volume contribution or curvature contribution: it is link to configurational entropy. It can be attributed to the difference in the degree of thermal expansion between the molecules.

The equation of state used in the PFP model is defined by the Eq. ( 9)

T v v v T v p 1 1 ~~3 1 3 1    (9)
ṽ v v     with       3 3 4 1 / 1 ~T T v i i i     
, where αi is the thermal expansion coefficient. The reduced temperature is determined considering the PFP EoS using Eq. [START_REF] Privat | Discussion around the paradigm of ideal mixtures with emphasis on the definition of the property changes on mixing[END_REF](11).

3 / 4 3 / 1 ~1 ~v v T   (10) T T v      1 3 1 ~3 / 1 (11)
The characteristic pressure can be calculated using Eq. ( 12)

2 * ~i Ti i i v T p    ( 12
)
where  is the thermal expansion coefficient and T  , the isothermal compressibility. The reduced configurational energy is determined by

i i v U / 1 ~  and the characteristic configurational energy is * * * i i i v p U 
. Concerning the heat capacity, we can estimate the reduced heat capacity at constant pressure following Eq. ( 13). The contact energy fraction in Eq. ( 14) is derived from hardcore volume fraction i  (Eq. ( 15)), surface fraction i  (Eq. ( 16)) and, molecular surface/volume ratios si.

* 2 2 * 1 1 * 1 1 2 1 1 p p p          (14) * 2 2 * 1 1 * 1 1 2 1 1 v x v x v x       (15) 2 2 1 1 2 2 2 s s s       (16)
The ratio s1/s2 can be approximated by the ratio of pure molar volume  

3 / 1 * 1 * 2 / v v
, leading to Eq. ( 17) for surface fraction.

3 / 1 * 1 * 2 2 1 2 1            v v    (17)
The reduced configurational energy for the mixture is

2 2 1 1 Ũ U U    
. The reduced heat capacity at constant pressure for the mixture is . Consequently, using the PFP equation of state and the different defined variables, the expression of the excess volume can be determined (Eq. ( 18)).

𝐻 𝐸 𝑥 1 𝑢 1 * +𝑥 𝑢 2 2 * = ( 1 𝑣 ̃+ 𝑇 4 3𝑣 ̃-1 3 ⁄ ⁄ -1 ) 𝛹 1 𝜃 2 𝑥 12 𝑝 1 * - ( 1 𝑣 ̃1 ⁄ -1 𝑣 ̃2 ⁄ ) 2 2 9 ⁄ 𝑣 ̃-2 3 ⁄ 4 3𝑣 ̃-1 3 ⁄ ⁄ -1 𝛹 1 𝛹 2 (18)
Unity of H E is imposed by the choice of unity of vi * and vi. Temperature is in Kelvin. The choice of unity of pressure (among other pi * ) imposed the unity of χ12.

We will use the parameters defined by Coquelet et al. [20]. We remind that for a pure component, the isothermal compressibility T i is related to the isentropic compressibility S via the Maxwell's relations T i =S + 2

v°iT/Cp

where vi [m 3 /mol] is the molar volume of the compound i, S is the isentropic compressibility (S =-1/v(v/p)S )

[Pa -1 ]. The molar volume, heat capacity, thermal expansion, isentropic compressibility and isothermal compressibility used [21] are calculated at atmospheric pressure using a second order polynomial function of temperature (Eq. ( 19)). The polynomial parameters a, b and c are reported in Table 10. The values of 12 parameters are presented in Table 12 for the furan + n-hexane and for the furan + 1-octanol binary systems. The Fig. 6 shows the results obtained for each temperature for the two binary systems. The different contribution to the molar excess enthalpies is also indicated. The results are also presented in supplementary files. We can conclude that the interactional interactions are predominant and confirm the selection of NRTL activity coefficient model used to correlate the data.

  c bT aT T F    2 (19)
For the two binary systems studied, and as it can be seen in Fig. 6 the main contribution to excess molar enthalpy is attributed to the interactional contribution. It highlights the importance of the 12 parameter in the data treatment and the dependence to the composition of this parameter.

Conclusion

The experimental excess molar volume and enthalpy data for 3 binary mixtures containing furan have been presented

at different temperatures and pressures, covering the entire composition range. The PFP model was used to correlate the data and identify the most predominant molecular interaction. The treatment confirms that the main molecular interactions are interactional type. Following the conclusion on our data treatment using PFP model, the selected NRTL activity coefficient model can be used to correlate experimental data. In order to understand the thermodynamic behaviour of oxygenated compounds, it could be of interest to represent the excess properties by a group contribution model. The thermodynamic model will make it possible to calculate the thermodynamic properties during the different steps of an industrial process such as manufacturing new products from the lignocellulosic material.

  K, 353.15 K and 383.15 K, using a vibrating tube densimeter DMA HPM from Anton Paar. The experiments were performed in a flow regime. The temperature inside the densimeter was controlled using a cryothermostat model F12-MA from Julabo, with an precision of 0.05 K. Its short-term temperature stability (over one experiment) was ±0.01 K while the long-term temperature fluctuation can reach 0.03 K. The pressure was maintained stable to 0.05 MPa using a back-pressure regulator from Swagelok (KPB1L0D412P200B0 model) and was measured by a Drück electronic (PMP4010 model) pressure gauge with a precision of 0.03 MPa.

Fig. 1

 1 Fig. 1 Schematic representation of the experimental setup for density measurements.

  K were realized at 0.5 and 1.0 MPa. The results show that the difference in molar excess enthalpy values is within experimental uncertainty. The densities and excess molar volumes of binary mixtures furan + n-hexane were obtained at 303.15, 353.15 and 383.15 K and 1.6 MPa. The excess enthalpies were measured at 313.15, 353.15 K and 383.15 K and, pressures 0.5 MPa and 1.0 MPa. Experimental data are reported in Table

Fig. 3

 3 Fig. 3 Furan + 1-octanol binary system; left: excess molar volumes at 1.65 MPa;  : 303.15 K;  : 353.15 K;  : 383.15 K. right: excess molar enthalpies at  : 313.15 K, 0.5 MPa;  : 353.15 K, 0.5 MPa;  : 383.15 K, 1.0 MPa. The full lines are RK correlations with 3 parameters. The dashed line is the RK correlation with 4 parameters.

Fig. 4

 4 Fig. 4 Furan + Toluene binary system; left: excess molar volumes at 1.65 MPa;  : 303.15 K;  : 353.15 K;  : 383.15 K. right: excess molar enthalpies at  : 313.15 K, 0.5 MPa;  : 353.15 K, 0.5 MPa;  : 383.15 K, 1.0 MPa.

Fig. 5

 5 Fig.[START_REF] Nala | Vapour-liquid equilibrium (VLE) for the systems furan+n-hexane and furan + toluene, measurements,data treatment and modeling using moleculat models[END_REF] Comparison of excess properties of binary systems with furan ( : 1-octanol;  : n-hexane;  : toluene); left: excess molar volumes of at 303.15 K and 1.65 MPa; Lines are RK correlations and right: excess molar enthalpies at 313.15 K and 0.5 MPa; The lines are NRTL correlations for furan + 1-octanol and furan + n-Hexane systems. RK fits are used for furan + toluene system.

.

  T*, p* and v* are the characteristics parameters obtained directly from the thermodynamic properties of the pure liquid. The tilde refers to an average reduced quantity. Concerning the volume

 12 Fig. 6

 126 Fig. 6 Excess molar enthalpy at 313.15 K (1), 353.15 K (2) and 383.15 K (3) of furan + n-hexane (A) and furan + 1-octanol (B) binary systems as a function of molar composition. Solid line: PFP theory, dotted line: Interactional contribution, dashed line: free volume contribution.

Table 1

 1 Chemicals used for experimental work

	Chemical name	Molecular Formula	CAS number	Supplier	purity	Molecular weight g.mol -1
	Furan	C4H4O	110-00-9	Acros Organics	99.0%	68.07
	n-hexane	CH3(CH2)4CH3	110-54-3	Acros Organics	99.0%	86.18
	Toluene	C6H5CH3	108-88-3	Fischer Scientific	99.0%	92.14

  3

Table 5

 5 Excess molar volume Redlisch-Kister parameters for binary mixtures at pressure p = 1.65 MPa

	T / K	A0	A1	A2	A3	SD*
			Furan -n-Hexane
	303.15	4.0805	0.2074	-0.3879		0.01
	353.15	5.5922	0.3302	-2.0069		0.01
	383.15	7.9925	-0.5985	0.3741		0.04
			Furan -Toluene
	303.15	-0.1738	0.1565	0.6465
	353.15	-0.3213	-0.02731	0.08688
	383.15	-0.2743	-0.1688	-0.2352
			Furan -1-Octanol
	303.15	1.3393	0.7241	-0.3206		0.02
	353.15	1.1539.10 -4	1.2499	-0.4081		0.03
	353.15	-0.01536	1.6555	-0.3386	-0.9965	0.01
	383.15	-0.9228	0.2099	0.5983		0.02
	𝑛 *Standard Deviation 𝑆𝐷 = √ ∑ (𝑉 𝑒𝑥𝑝 𝐸 -𝑉 𝐶𝑎𝑙𝑐 𝐸 ) 2 𝑛

Table 6

 6 Excess molar enthalpies Redlisch-Kister parameters for binary mixtures with pressure between 0.1 and 1 MPa.

	T / K	A1	A2	A3	SD**
			Furan -n-Hexane
	303.15*	4815.70	847.970	514.966	7.2
	313.15	4754.60	706.468	74.2019	6.0
	353.15	4357.94	763.296	83.1453	8.5
	383.15	3813.27	725.836	37.6165	6.4
			Furan -Toluene
	313.15	219.235	-31.961	51.1446	2
	353.15	195.717	-58.779	6.17613	2
	383.15	153.245	4.4940	-2.08364	4
			Furan -1-Octanol
	313.15	5870.00	3718.04	3098.19	48
	353.15	7060.40	3603.94	2539.81	26
	383.15	6573.37	2683.14	975.98	35
	𝑛 *RK correlation for literature values [Guillen et al. [9]]; ***Standard Deviation 𝑆𝐷 = √ ∑ (𝐻 𝑒𝑥𝑝 𝐸 -𝐻 𝐶𝑎𝑙𝑐 𝐸 ) 2 𝑛

  N is the number of data points, subscript cal corresponds to calculated value and, exp corresponds experimental value.

	T / K	a12	a21	*AAD (%)
			Furan + n-hexane
	313.15 K	1005.18	455.96345	0.8
	353.15 K	1030.26	287.06292	0.8
	383.15 K	974.58	149.93793	0.7
			Furan + 1-octanol
	313.15 K	2922.15	835.402577	2
	353.15 K	2858.46	812.992373	0.7
	383.15 K	2206.00	460.572079	1.4
	*𝑨𝑨𝑫 = (𝟏𝟎𝟎/𝑵) ∑|(𝐻 𝑐𝑎𝑙 𝐸 -𝐻 𝑒𝑥𝑝 𝐸	)/𝐻 𝑒𝑥𝑝 𝐸	|,

Table 10

 10 Values of the second order polynomial expression (Eq. (18)) for pure compounds.The PFP model was applied for the 2 binary systems at 313.15, 353.15 and 383.15K. In order to adjust the 12 parameter, we have minimized the objective function (Eq. 20) using PFP parameters derived from Eq. (21), for each pure component property. Table11presents the values of the parameters of each pure component at the corresponding temperature.

	Property F	a	b	c
		1-Octanol
	v/ cm 3 .mol -1	1.70.10 -04	0.032	133.8
	 /K -1	-2.12.10 -09	2.695.10 -06	2.3.10 -04
	Cp /J.mol -1 .K -1		1.070	-14.51
	S  /Pa -1	1.56.10 -14	-5.3.10 -12	8.7.10 -10

Table 11

 11 Physical constants of pure compounds at different temperatures, used for calculations of excess molar volume V E with PFP model: vi°: molar volume, i coefficient of thermal expansion T isothermal compressibility, reduction parameters of volume vi*, energy Ui* and pressure pi*.

	Compound	vi° / cm 3 .mol -	10 3 i / K -1	10 3 T /	vi* /	pi* / MPa Ui* / kJ.mol -1
		1		MPa -1	cm 3 .mol -1
				T = 313.15 K
	Furan	74.776	1.549	1.456	54.840	619.59	33.978
	n-hexane	134.324	1.457	1.968	99.708	420.67	41.944
	1-Octanol	160.540	0.865	0.851	130.660	480.51	62.783
				T = 353.15 K
	Furan	79.821	1.697	2.100	56.073	578.44	32.435
	n-hexane	142.838	1.610	2.843	101.442	396.59	40.231
	1-Octanol	166.358	0.917	1.074	131.496	482.76	63.482
				T = 383.15 K
	Furan	84.123	1.783	2.688	57.490	544.24	31.288
	n-hexane	150.124	1.702	3.650	103.607	375.125	38.865
	1-Octanol	171.078	0.951	1.273	132.448	477.61	63.258

Table 12 χ12

 12 parameters used in the PFP theory for binary systemsWe have considered the same temperature dependency of the cross parameter 12 (Eq.21) like inCoquelet et al. [20].

	T /K	0 12 	a	b
		furan + 1-octanol
	313.15	80.76	1.2595	0.9537


G Brillant

J Martin
email: jimmy.martin@irsn.fr

B Fourré

EXPERIMENTAL STUDY OF AN ADIABATIC TWO-PHASE FLOW IN A SIMULATED PWR SPENT FUEL BUNDLE

Keywords: Spent Fuel Pool, loss-of-cooling/loss-of-coolant accident, DENOPI project, Two-phase flow

Within the framework of the post-Fukushima actions, an experimental project named DENOPI has been launched with the aim to study the behavior of spent fuel pools under loss of cooling and loss of coolant conditions. Within this project, the MEDEA facility is intended to study the thermalhydraulics of a fully covered or uncovered spent fuel bundle. It consists in a hydraulic loop connected to an experimental test section in which a simulated one meter high PWR spent fuel bundle is located. The MEDEA-overflow experiments, presented in this study, consist in adiabatic air/water flows in a fully covered unheated bundle. Axial pressure and void fraction distributions were measured using pressure transducers. Overall, the tests exhibited two different flow regimes. Indeed, at low gas superficial velocities, a bubbly flow set up within the bundle whereas a transition to slug flow was obtained at larger ones.

INTRODUCTION

The vulnerability of Spent Fuel Pools (SFP) has been replaced to the front of the stage since the Fukushima-Daiichi accident [START_REF] Wang | Study of Fukushima Daiichi nuclear power station unit 4 spent-fuel pool[END_REF][START_REF] Wu | Analysis of accidental loss of pool coolant due to leakage in a PWR SFP[END_REF][START_REF] Chen | Accident mitigation for spent fuel storage in the upper pool of a Mark III containment[END_REF][START_REF] Song | Severe accident issues raised by the Fukushima accident and improvements suggested[END_REF] and several kinds of actions have been launched in order to understand and reduce the risk and consequences of potential accidents on SFP. For instance, a working group set up at OECD/NEA issued a status report on SFP Loss-of-Cooling and Loss-of-Coolant Accident (SFP-LOCA), that aims at providing a summary of the status of SFP accident and mitigation strategies, a brief review of the state of the art of the simulation tools potentialities for SFP-LOCA assessment and a proposal for some additional research actions [5]. Moreover, a European NUGENIA project named Air-SFP was performed in 2015-2016, in order to assess the uncertainties of Severe Accident (SA) codes in dealing with SFP-LOCA and to identify needs of modeling improvement [START_REF] Coindreau | Severe accident codeto-code comparison for two accident scenarios in a spent fuel pool[END_REF]. Recently, a Phenomena Identification Ranking Table (PIRT) activity related to SFP-LOCA was carried out at OECD/NEA [START_REF]PIRT: R&D priorities for loss-of-cooling and loss-of-coolant accidents in spent nuclear fuel pools[END_REF]. From these activities, it appears that the validation database of most computer codes currently used for SFP deterministic safety analysis (system codes developed on the basis of in-reactor incidental or accidental conditions) needs to be extended to SFP configurations. Indeed, the thermal-hydraulics of a SFP is mainly based on natural convection flows (gaseous, liquid or two-phase flows), developed at atmospheric pressure with a relatively low heat load, compared to in-reactor conditions. These current codes haven't been primarily developed for natural convection flows and their applicability to SFP has to be improved.

In this context, the DENOPI project [START_REF] Mutelle | A new research program on accidents in spent fuel pools: the DENOPI project[END_REF][START_REF] Martin | Overview of the IRSN DENOPI project on spent fuel pool in loss-of-cooling and loss-ofcoolant accident conditions[END_REF][START_REF] Martin | The IRSN DENOPI project: a research program on spent-fuel-pool loss-of-cooling and loss-of-coolant accidents[END_REF] has been launched by the IRSN (the French Technical and Scientific Support Organization) in collaboration with several research laboratories with the aim of studying the behavior of spent fuel pools under Loss-of-Cooling and Loss-of-Coolant Accidents (SFP-LOCA). The DENOPI project is a research program including both experimental and modeling activities. Its purpose is to provide code developers with an experimental database, made up of SFP integral and separate effect tests, for code improvement and validation. The project is divided into three axes, each corresponding to a specific spatial scale involved in a SFP-LOCA. The first axis is related to the twophase natural convection flows occurring at the pool scale, prior to the fuel uncovery. The thermalhydraulics of a typical PWR (Pressurized Water Reactor) fuel bundle prior and after the fuel uncovery is investigated in the second axis. In particular, the efficiency of spray cooling systems as a mitigation measure is assessed. Finally, the third axis is dedicated to the fuel cladding degradation by steam-air mixture oxidation after fuel uncovery.

The MEDEA mock-up was built within the framework of the DENOPI project, and more specifically in its second axis. The main component of this facility is a one meter high unheated rod bundle. The objective of this experimental set-up is to get new insights on the physical phenomena involved during uncovering and water spraying of a fuel bundle stored in a spent fuel pool. The MEDEA program has two main steps. The first one is focused on air/water experiments, and the second one on steam/water experiments. The air/water experiments are of two kinds: the study of flooding in case of water spraying on a completely uncovered bundle (MEDEA-flooding tests) and the study of the void distribution along a fully covered assembly in an air/water co-current flow (MEDEA-overflow tests). The results of the MEDEA-overflow tests are presented in this article. The MEDEA-overflow tests aim at studying the pressure loss and void fraction along a PWR fuel bundle for a co-current air/water flow at room temperature and atmospheric pressure. These measurements are required for the validation of thermalhydraulics numerical tools, especially regarding the drift flux models for the low pressure domain of natural flow in a fuel bundle. This air/water flow study is a first approach to observe and measure the flow pattern and void fraction for a co-current gas/liquid flow awaiting the ASPIC facility. Actually, the ASPIC facility, currently under construction, is a full height 17x17 assembly with heated rods (80kW, up to 600 °C) that will enable several kinds of experiments and scenarios for SFP accidents at the assembly scale.

This article is organized as follows. In section 2, a presentation of the MEDEA device and measurement apparatus is driven. The test parameters and the experimental protocol are laid out in section 3. Then, the results of the test series are stated and analyzed in section 4. Finally, the main outcomes of this experimental study are drawn in the conclusion.

DESCRIPTION OF THE OVERFLOW CONFIGURATION OF THE MEDEA FACILITY

Several experimental rigs are planned in the framework of the DENOPI project. The MIDI facility [START_REF] Martin | Overview of the IRSN DENOPI project on spent fuel pool in loss-of-cooling and loss-ofcoolant accident conditions[END_REF][START_REF] Trégourès | The DENOPI project: a research program on SFP under loss-of-cooling and loss-ofcoolant accident conditions[END_REF] will be devoted to the investigations on large scale flow patterns and onset of boiling in SFP before the dewatering of the fuel bundles. The MEDEA and ASPIC test rigs are dedicated to the study of the physical phenomena involved at the assembly scale during a SFP deflooding. All these new facilities will belong to the IRSN platform THEMA (THErmalhydraulics for Mitigation of Accidents). The overflow configuration of the MEDEA facility is composed of a test section with a one meter high rod bundle, an air injection line connected to the bottom of the test section, a water injection line connected to the bottom of the test section, and a water gathering circuit for the overflow liquid (Cf. figure 1).

The MEDEA assembly consists in a 17x17 rod bundle with a reduced height of 1240 mm for a weight of about 80 kg. The rods have an outer diameter of 9.5 mm and the pitch of the assembly is 12.6 mm. The 25 guide tubes of the bundle have an outer diameter of 12 mm. The three grids of the bundle consist in a mixing grid in between two support grids. There is 530 mm between the upper support grid and the mixing grid and 400 mm between the mixing grid and the lower support grid. The central test section, in which the assembly is inserted, is a square tube with an internal dimension of 225 mm corresponding to a cell of a spent fuel pool rack. The top nozzle, spacer grids, and rods used to build the assembly are representative to a typical PWR rod bundle. The free flow area of the top nozzle is about 205 cm 2 . Below the test section with the rod bundle, the lower section aims at injecting the water flow and the air bubbles and obtaining a stationary flow at the bottom of the rod bundle. The length of the lower section is 1416 mm (about 6 hydraulic diameters). Above the test section, the upper head is connected to the gathering water line that collects the overflow. Note that about 317 mm below the overflow level could be placed a Wire Mesh Sensor (WMS) [START_REF] Prasser | A new electrode-mesh tomograph for gas-liquid flows[END_REF][START_REF] Schleicher | Refined reconstruction of liquid-gas interface structures for stratified two-phase flow using wire-mesh sensor[END_REF] which provides measurements of the void fraction on a grid pattern. These measurements are under analysis and will be presented and discussed in a forthcoming article.

In order to regulate the air volume flowrate injected in the test section, a flowmeter/valve group is incorporated in the air injection line. This flowmeter/valve group is used in automatic mode. That way, the operator sets the target flowrate which is stabilized by means of a PID controller. The water injection line is connected to a tank with a capacity of 2000 l. The water is pumped from that tank and flows through a flowmeter/valve group. This group can either be used in manual or automatic mode to set up the water mass flowrate in injection line. The water flows through the central section and the rod bundle. Note that a bypass line can be partially opened with manual valve to send back part of the water flow directly to the main tank. The water flow above the rod bundle leads to an overflow that is collected and sent back by gravity to the main tank by means of the water gathering line.

TEST PARAMETERS AND EXPERIMENTAL PROTOCOL

The MEDEA-overflow test series has three main parameters: the water mass flowrate, the air volume flowrate and the geometrical configuration. Two configurations have been considered: with (configuration C1) and without (configuration C2) the rod bundle inside the central test section. That way, the impact of the rod bundle on the evolution of the pressure loss and void fraction all along the test section can be directly evaluated. The range of the water mass flowrate is 400 g/s to 2400 g/s and the range of the air flowrate is 10 Nl/min to 180 Nl/min (flowrate at 0 °C and 1 atm).

The experimental protocol of the MEDEA-overflow tests is based on the variation of the air flowrate injection while the water mass flowrate and the geometrical configuration are unchanged. Actually, the series started with a test at zero air flowrate to determine the default signal of all the pressure transducers. That way, the transducers inherent offset (e.g. due to transducer orientation) as well as the potential dynamic pressure offset can be subtracted from the other tests with positive gas flowrates. Note that the free flow section is not constant along the test section which leads to variations of the water velocity and thus of the dynamic pressure. For each test, the signal of the pressure transducers is recorded over at least 10 min at 1 Hz. Hence, stabilized mean and standard deviation values of the pressures can be calculated. The locations of all the pressure transducers are reported in figure 1. The uncertainties of the pressure measurements, considering both the sensor and the supply chain, are 0.05 mbar for the differential pressure transducers ݀ܲ ܣ -݀ܲ ܩ and 1.5 mbar for the absolute pressure transducers ܲ ݑ and ܲ ݀ . As far as the flowrate measurements are concerned, the uncertainties are 1 % for the air flow and 0.1 % for the water flow. Regarding the absolute pressure measurements, the pressure difference ܲ ݀ -ܲ ݑ can be written as (terms with air density are neglected as well as the dynamic pressure drop between two sensors):

݀ܲ ܼ = ܲ ݀ -ܲ ݑ = (1 -ߙ ܼ ) • ߩ ݓ • ݃ • ℎ + ߩ ݓ • ݃ • ܪ( -ℎ) (1)
Where ܪ is the elevation difference between the two transducers, ℎ the elevation difference between the two connection holes on the central section, and ߙ ܼ the mean void fraction along the test section between the two holes where the two transducers ܲ ݑ and ܲ ݀ are connected. Therefore, the void fraction ߙ ܼ can be estimated as:

ߙ ܼ = (ߩ ݓ ܲ݀-ܪ•݃• ܼ ) ߩ ݓ •݃•ℎ (2)
Concerning the differential pressure measurements, the following expression can be written:

݀ܲ ݅ -݀ܲ ݅ 0 = ߩ ݓ • ݃ • ݀ℎ ݅ -(1 -ߙ ݅ ) • ߩ ݓ • ݃ • ݀ℎ ݅ ; ݅= ,ܣ … , ܩ (3)
with ݀ܲ ݅ 0 the offset signal of the differential pressure transducers and dh the elevation difference between the two connections of the dedicated transducer. Therefore, the void fraction can be calculated as:

ߙ ݅ = (݀ܲ ݅ -݀ܲ ݅ బ ) ߩ ݓ •݃•݀ℎ ݅ ; ݅= ,ܣ … , ܩ (4)
The void fractions profiles presented hereafter were calculated by means of these expressions.

RESULTS

Pressure measurements for the series at a water flowrate of 1200 g/s

For this test series, eleven values of the air volume flowrate are considered in the range 0 Nl/min to 180 Nl/min and the water mass flowrate is kept constant to the value of 1200 g/s during each test (meaning each plateau of air flowrate). As the friction inside the test section evolves with the air flowrate, the opening rate of the valve on the water line was adapted on each test to stick to the target value of the water mass flowrate. In doing so, the mean value of the water mass flowrate is kept in the range 1198 g/s to 1223 g/s (2 % maximum deviation). It can be noticed that, for all tests, the standard deviation of the water mass flowrate is around 1.5 g/s.

The mean value of the pressure difference ܲ ݀ െ ܲ ݑ on the test section is given in figure 2a. This pressure difference decreases from 196.0 mbar to 165.9 mbar while the air flowrate rises from 0 Nl/min to 180 Nl/min. The measurements from the differential pressure transducers, gathered in figure 2b, reveal a variation in the range 0 mbar to 6 mbar. Concerning the standard deviation of the pressure signals (Cf. figure 3), it can be noticed that the standard deviation of the differential pressure transducers clearly increases with the air volume flowrate. However, this trend cannot be observed in figure 3a for the absolute pressure measurement on the full height of the section. This can probably be imputed to the higher uncertainty of the absolute pressure transducers. Besides, it can be noted that the standard deviation above the rod bundle (curve A, figure 3b) is higher than the standard deviation in the assembly area, revealing a more perturbed flow in this region.

Void fraction with/without assembly at Q w = 1200 g/s

The void fractions, calculated by means of the pressure measurements and equations ( 2) and ( 4), are plotted in figure 4a for the test series with the rod bundle inserted into the test section. Firstly, the lower values of the void fraction are measured above the rod bundle (transducer A). In fact, these smaller values may be ascribed to the higher section radius above the assembly and to the recirculations observed below the overflow water level. It can be noticed a stabilization at this location of the void fraction around 7 % for the air volume flowrates higher than 120 Nl/min. The highest values of the void fraction are measured just below the rod bundle (transducer F) and are due to a plug created by the lower plate below the assembly. Note that this plug grows, as well as the void fraction, as the air volume flowrate rises. Intermediary void fraction are observed all along the rod bundle (transducers B-E) as well as in the lower part of the test section (transducer G). The impact of the grids on the void fraction profiles can be hardly explained. In fact, the transducers B (with a support grid) and D (with a mixing grid) lead to similar void fraction profiles but a different behavior is noted for the transducer E (with a support grid). The void fraction in the lower part of the test section (transducer G) is lower than the void fraction along the rod bundle. This observation can be imputed to the higher friction in the assembly (due to both rods and grids). The mean void fraction (ߙ ܼ ) over the full height of the test section measured by the two absolute pressure transducers is among the other void fraction curves measured locally by all the differential pressure transducers. This test series was repeated without the rod bundle in the central test section in order to give information for the code development and validation. The void fraction measured during the test series without the assembly in the central test section are reported on figure 4b as a function of the air volume flowrate. It can be noticed that the measurements for the transducers B to E are quite clustered due to the absence of the assembly. The void fraction differences between the tests with and without the subassembly in the central section are given on figure 5. The differences of void fraction between with and without bundles tests increase with the air volume flowrate. It is consistent with the fact that the fluid frictions rises with the velocity and hence with the flowrates. However, the differences remain below 5 % for the range of value of the air and water flowrates. It can be observed that the larger differences are observed for the transducers B to E which corresponds to the area of the rod bundle.

Repeatability of the test series

The repeatability of the measurements was checked on the test series at Q w = 1200 g/s with the rod bundle inserted into the test section. The mean and standard deviation values of the water mass flowrate and the air volume flowrate are similar in both test series and have a negligible impact on further measurements. Regarding the absolute pressure, a difference from 1 mbar to 5 mbar is noticed which leads to discrepancies up to 3 % of the void fraction (Cf. figure 6). The same trend is observed on the differential pressure curves with a difference up to 2 mbar from one test to the other that induces a variation on the void fraction up to 30 % (meaning 7 % void fraction absolute difference). These discrepancies put in light the intrinsic highly unstationary behavior of the hydraulics phenomena involved in the test section.

Analysis of the whole test series

Overall, the two-phase flows simulated during the tests exhibited two differing behaviors, depending on the investigated geometrical configuration. Those behaviors are discussed below.

First of all, with no rod bundle in the test section, the two-phase flows always correspond to a bubbly flow, with no noticeable bubble coalescence along the vertical axis. In this configuration, any increase of the gas superficial velocity leads to a gentle variation of the void fraction in the duct, as seen in figure 7b. At low gas superficial velocity, one may note that no clear trend can be observed when the liquid superficial velocity is varied. Visually, the flow exhibits a high level of randomness with bubbles fluctuating in all directions, in spite of their upward net movement. The two-phase flows within a rod bundle exhibit a different behavior. In such a case, a bubbly flow is first visualized for the gas superficial velocities lower than 5 cm/s. When more gas is added to the test section, the flow organizes differently and a cap-bubbly flow is reached, in which big gas pockets can be identified. The latter results from the coalescence of individual bubbles, promoted by the confinement imposed to the flow by the presence of the rod bundle. From that point, some big gas pockets accumulate sometimes below the rod bundle upper tie plate, thereby leading to macro-scale flow recirculations. Those recirculations are evidenced by tiny bubbles flowing downwardly and counter-currently to the main flow direction. As seen in figure 7a, an increase of the gas superficial velocity also leads to bigger void fractions. But, contrarily to the case with no bundle, a clear trend is observed here regarding changes in liquid superficial velocities: when more liquid is added to the test section, the void fraction reduces accordingly. Noticeably, for high liquid superficial velocities, this latter trend yields void fractions very close to the ones obtained with no rod bundle. We assume that an increasing liquid superficial velocity promotes bubbles dispersal by turbulent mixing within the test section and prevents big gas pockets formation and accumulation below the upper tie plate: qualitative flow observations show that those pockets are hardly visible in the test section at higher liquid superficial velocities.

CONCLUSIONS

One objective of the DENOPI project is to improve our understanding of the physical phenomena involved at the fuel assembly scale in a spent fuel pool in case of loss of cooling or loss of coolant accidents. The MEDEA-overflow experiments consist in a co-current air/water flow through a one meter height unheated rod bundle inserted into a SFP rack. In these experiments, no clear impact of the mixing grids and support grids on the void fraction profiles could be evidenced. Besides, two flow regimes were put in light. At low gas superficial velocities, a bubbly flow set up within the bundle whereas a transition to slug flow was observed at larger velocities. This experimental data will be considered in the near future to improve the modeling of this kind of flow by system codes especially by determining the most adequate drift flux model to recommend. Furthermore, the ASPIC facility, a heated full height assembly currently under construction, will enable in the near future several kinds of experiments and scenarios for SFP accidents at the assembly scale. For instance, boil-off scenarios will be considered with the evaluation of the efficiency of a water spray to cool down a partially/totally dry fuel bundle. Besides, measures of the pressure loss and void fraction along the assembly will be undertaken with a steam/water flow and compared to the results of the present study.

Figure 1 .

 1 Figure 1. Locations of the pressure measuring holes (A to G) on the MEDEA-overflow test section.

Figure 2 .

 2 Figure 2. Mean pressure measurements for the tests at a water flowrate of 1200 g/s.

Figure 3 .

 3 Figure 3. Standard deviation of the pressure measurements for the tests at a water flowrate of 1200 g/s.

Figure 4 .

 4 Figure 4. Void fraction for the tests at a water mass flowrate of 1200 g/s.

Figure 5 .

 5 Figure 5. Void fraction difference for the tests with/without assembly at a water flowrate of 1200 g/s.

Figure 6 .

 6 Figure 6. Comparison of the void fraction for the two tests series at a water flowrate of 1200 g/s (repeatability).

Figure 7 .

 7 Figure 7. Mean void fraction over vertical direction as a function of the gas superficial velocity (J g ).

  gas velocity (m/s) ݀ܲ ܣ pressure difference of the differential pressure transducer at location A (Pa) ܲ ݀ pressure of the absolute pressure transducer at location do (Pa) Q w water mass flowrate (kg/s) Q air air volume flowrate (Nl/min) ߙ ܣ void fraction at location A (-) ߩ ݓ water volumetric mass density (kg/m 3 )

ACKNOWLEDGMENTS

The DENOPI project is part of the "Investment for the future" program funded by the French Government within the framework of the post-Fukushima surveys identified as major safety issues (contract number ANR 11 -RSNR 006).