Minh Pham

Doan

Nzihou Ange

Sharrock Patrick

Production of Sodium Dihydrogen Phosphate Using Sodium Chloride and Orthophosphoric Acid

Up-to-date, standard synthesis processes of sodium phosphates are based on the neutralization of orthophosphoric acid by sodium hydroxide or sodium carbonate. To the best of our knowledge, the reaction between sodium chloride and orthophosphoric acid has not yet been reported in the literature. This study demonstrated the feasibility of the synthesis of sodium dihydrogen phosphate by the reaction of sodium chloride with orthophosphoric acid. A detailed parametric study was done, wherein the influence of reaction temperature, air flow rate, water addition flow rate, and reaction time was investigated. As expected, the evacuation of hydrogen chloride (HCl) was crucial for getting high conversion of NaCl to HCl. The highest conversion of NaCl to HCl reached 91% after 24 h of synthesis reaction. Sodium dihydrogen phosphate (NaH 2 PO 4 ) was the main product. Pyrophosphate (P 2 O 7 4-) was only formed at trace amounts, and any trace of trimetaphosphate (P 3 O 9 3-) was observed.

INTRODUCTION

Phosphates include a large number of chemical products and can be divided into different categories: orthophosphates, linear polyphosphates, cyclic metaphosphates, and cross-linked phosphates. They are widely used in various applications such as fertilizer, [START_REF] Walan | Phosphate rock production and depletion: Regional disaggregatedmodeling and global implications[END_REF][START_REF] Ahmad | Parametric evaluation using mechanistic model for release rate of phosphate ions from chitosancoated phosphorus fertiliser pellets[END_REF][START_REF] Kulakovskaya | Inorganic polyphosphate in industry, agriculture and medicine: Modern state and outlook[END_REF] food industries, [START_REF] Kulakovskaya | Inorganic polyphosphate in industry, agriculture and medicine: Modern state and outlook[END_REF][START_REF] Capuano | Comparison of a sodium-based and a chloride-based approach for the determination of sodium chloride content of processed foods in the Netherlands[END_REF] energy, [START_REF] Kulakovskaya | Inorganic polyphosphate in industry, agriculture and medicine: Modern state and outlook[END_REF][START_REF] Grolleau | Calendar aging of commercial graphite/LiFePO 4 cell-Predicting capacity fade under time dependent storage conditions[END_REF][START_REF] Zhang | Structure and performance of LiFePO 4 cathode materials: A review[END_REF] environment, [START_REF] Pham Minh | Hydroxyapatite starting from calcium carbonate and orthophosphoric acid: synthesis, characterization, and applications[END_REF][START_REF] Pham Minh | Calcium phosphate based materials starting from calcium carbonate and orthophosphoric acid for the removal of lead(II) from an aqueous solution[END_REF][START_REF] Pham Minh | Apatitic calcium phosphates: Synthesis, characterization and reactivity in the removal of lead(II) from aqueous solution[END_REF][START_REF] Pham Minh | Hydroxyapatite gel for the improved removal of Pb 2+ ions from aqueous solution[END_REF][START_REF] Raii | Lead and cadmium removal from aqueous solution using an industrial gypsum by-product[END_REF][START_REF] Kemiha | Highly porous calcium hydroxyapatite-based composites for air pollution control[END_REF] catalysis, [START_REF] Vedrine | Molecular description of active sites in oxidation reactions: Acid-base and redox properties, and role of water[END_REF][START_REF] Didenko | Partial catalytic oxidation and condensation of methane by oxygen and Sulphur[END_REF][START_REF] Boukha | Methane dry reforming on Ni loaded hydroxyapatite and fluoroapatite[END_REF] etc. Orthophosphoric acid is industrially produced from the dissolution of phosphate rocks by sulfuric acid. [START_REF] Gobbitt | Yara hemihydrate (HH) and hemidihydrate (HDH) processes for phosphoric acid production[END_REF][START_REF] Azaroual | Thermo-kinetic and physico-chemical modeling of processes generating scaling problems in phosphoric acid and fertilizers production industries[END_REF] The neutralization of orthophosphoric acid with a base forms the corresponding orthophosphate salts. On the other hand, most pyrophosphates and polyphosphates are generally obtained by the condensation of orthophosphates. For example, sodium trimetaphosphate (STM, Na 3 P 3 O 9 ) is produced on an industrial scale by heating sodium dihydrogen phosphate (NaH 2 PO 4 ) at 500 °C for 5 h. [START_REF] Pham Minh | A new route for the synthesis of alkali polyphosphate from economical starting materials: preparation and characterization of sodium cyclotriphosphate[END_REF][START_REF] Pham Minh | A new route for the synthesis of alkali polyphosphate from economical starting materials: Part IIInfluence of reaction conditions[END_REF][START_REF] Pham Minh | Onestep synthesis of sodium trimetaphosphate (Na 3 P 3 O 9 ) from sodium chloride and orthophosphoric acid[END_REF] Among the large number of phosphate compounds, alkali orthophosphates appear as the most important ones, in particular for being used as fertilizers and food additives. [START_REF] Kulakovskaya | Inorganic polyphosphate in industry, agriculture and medicine: Modern state and outlook[END_REF] As mentioned above, alkali orthophosphates are obtained by the neutralization of orthophosphoric acid and the corresponding alkali hydroxide, such as NaOH. NaOH is industrially produced by the electrolysis of brine, the aqueous solution of sodium chloride. So, it is of interest to get sodium orthophosphates and sodium phosphates, directly from orthophosphoric acid and sodium chloride, as shown in eq 3.

Classical production of sodium dihydrogen phosphate: (2)

+ → + + Δ °+ - H 2NaCl (
The studied process in this work: (3) However, to the best of our knowledge, there has not yet been an industrial application of the direct reaction between orthophosphoric acid and sodium chloride. This seems to be the consequence of poor literature data on this reaction, with only some rare work at the laboratory scale reported. [START_REF] Blumberg | Process for the manufacture of mono-alkali metal phosphates[END_REF]22 The first one described the synthesis of monoalkali or monoammonium orthophosphates (MH 2 PO 4 , with M: K, Na, and NH 4 ) from the reaction of MCl with H 3 PO 4 in the presence of an organic solvent such as n-octane, white sprit, kerosene, tetrachloroethane, etc. [START_REF] Blumberg | Process for the manufacture of mono-alkali metal phosphates[END_REF] The reaction was carried out at 12-160 °C in order to form the HCl-H 2 O azeotrope. The role of the organic solvent is probably for facilitating the separation of the HCl-H 2 O azeotrope from the reaction mixture. Experiments have been done with both batch and continuous reactors which showed interesting results. The main drawbacks of this process were the presence of several steps (multisteps synthesis) and the use of an organic solvent.

+ → + Δ °- - H NaCl(
Zhou et al. 22 described the synthesis of NaH 2 PO 4 and Na 2 HPO 4 from NaCl and H 3 PO 4 using an organic amine. This amine is not miscible with the aqueous phase and is for trapping HCl formed from the reaction of NaCl and H 3 PO 4 . Thus, the organic phase contains HCl while the aqueous phase contains NaH 2 PO 4 or Na 2 HPO 4 . These two phases are easily separated by simple decantation. Crystals of NaH 2 PO 4 or Na 2 HPO 4 are obtained by cooling down the aqueous fraction. On the other hand, the organic solvent is regenerated by adding of ammonium hydroxide, which forms NH 4 Cl as a byproduct. By varying the organic amine, Na 2 HPO 4 could be obtained, indicating that the process was performed for the conversion of the second proton of orthophosphoric acid. However, the main drawbacks of this process relate to the use of an organic amine, as well as the consumption of a non-negligible quantity of NH 4 OH.

In this work, we demonstrated the feasibility of the synthesis of NaH 2 PO 4 from NaCl and H 3 PO 4 by a one-step synthesis process, using only water as a solvent.

MATERIALS AND METHODS

Sodium chloride powder with the volume-mean diameter (d NaCl ) of 375 μm from Acros Organics and 85 wt % orthophosphoric acid from Merck were used. Both the reactants were pure analytical grade. Distilled water was also used for the synthesis reaction.

Three different glass reactors were used in this work (Figure 1). The first one (N°1) was a 250 mL three-neck round-bottom flask connected with a vapor condenser. The second reactor (N°2) was similar to the first one but without a vapor condenser. The third reactor (N°3) was a 250 mL erlenmeyer flask modified by adding two identical necks.

The heating of the reaction medium was realized via a bath of thermal oil, which can be heated up to 300 °C. The oil bath was heated by a heating plate, with the regulation of the temperature by thermocouple 2 inserted in the oil bath. For the reactor N°3, the temperature of the reaction mixture was monitored by thermocouple 1.

For a given synthesis reaction, 20.45 g (or 0.35 mol) of fine powder NaCl and 40.35 g of 85 wt % H 3 PO 4 (or 0.35 mol H 3 PO 4 and 6.05 g H 2 O) were introduced into the reactor under stirring (500 rpm). This corresponded to the initial molar ratio of sodium to orthophosphate equal to 1/1. After heating the oil bath to the desired temperature between 80 and 130 °C, the glass reactor was immersed in this oil bath for starting the reaction. During the reaction, liquid water could be added with a controlled flow rate using a peristaltic pump through a neck of the reactors, as illustrated in Figure 1 (reactor N°3). An air flux passed over the surface of the reaction mixture, in order to evacuate HCl-vapor mixture out of the reactor. This mixture was trapped into a water washing flask to prevent air pollution. The quantity of Cl -in the trap was analyzed by ion chromatography [Dionex IonPac AS 19 (4 × 250 mm) column, conductivity detector]. This allows for calculating the conversion of NaCl to HCl by the following equation:

= × X n n 100% f i NaCl Cl Cl ( 4
)
with n Cl i the initial quantity of NaCl introduced into the reactor, and n Cl f the quantity of Cl evacuated to the trap. At the end of the synthesis reaction, the final product in the reactor was completely dissolved and diluted with water before analysis by ion chromatography and by ICP-AES (HORIBA Jobin Yvon Ultima 2), in order to make mass balance. In this work, we have studied the influence of the following parameters on the conversion of synthesis: (1) the reaction temperature, (2) the air flow rate, (3) the water addition flow rate, (4) the reaction time, and ( 5) the vapor condensation.

RESULTS AND DISCUSSION

3.1. Reactor Validation. As mentioned above, three glass reactors were used in this work (Figure 1). The first tests were carried out under the same reaction conditions (80 °C), in order to choose the highest performing reactor. The results are presented in Figure 2. The reactor N°3 showed the highest efficiency with 7% of NaCl conversion after 4 h of reaction. The reactor N°2 had much lower efficiency compared to the reactor N°3 with only 2% of conversion after 4 h of reaction and did not progress between 3 and 4 h of reaction. The reactor N°1 had the lowest efficiency with only trace amounts of HCl formed after 4 h of reaction.

HCl was formed according to eq 3 and needed to be evacuated out of the reaction mixture to favor the formation of the desired product. For the reactor N°1, HCl and vapor were condensed and returned to the reaction medium. So, only trace amounts of HCl was found in the trap using this reactor. For the reaction N°2, HCl could get out of the reactor but the section of the neck for HCl outlet was small compared to the reactor walls. This made the evacuation of HCl difficult. We observed the formation of aqueous HCl drops on the reactor walls which returned to the reaction mixture during the reaction with the reactor N°2. On the other hand, the reactor N°3 favored the evaporation of HCl-vapor mixture thanks to its large outlet. So, the reactor N°3 was thereafter used for other studies.

3.2. Influence of the Reaction Temperature. The temperature of the oil bath was controlled using the thermocouple 2 (Figure 1). On the other hand, the temperature of the reaction mixture was monitored by the thermocouple 1. Figure 3 shows the influence of the reaction temperature on the conversion of NaCl to HCl. When the oil bath was kept at 80 °C, we observed that the temperature of the reaction mixture reached 69-71 °C, leading to low efficiency of the process (less than 10% of conversion after 4 h of reaction time). The increase of the temperature of the oil bath to 100, 115, and 130 °C allowed increasing the performance of the process. The highest performance was found for 115 °C with 54% of conversion after 7 h of reaction time. In fact, when the temperature of the oil bath was set at 115 °C, the temperature of the reaction mixture varied in the range of 108-126 °C, which is close to the boiling temperature of the HCl-H 2 O azeotrope (108.6 °C). So this facilitates the elimination of HCl from the reaction mixture. A thermodynamic calculation of Gibbs energy minimization can help to suggest how the azeotrope can be best used to get HCl out of the reactor. However, the water content in the reaction mixture varied at a given reaction time. So, we could not calculate Gibbs energy minimization for such a system. The temperature of the oil bath of 130 °C was slightly less efficient than 115 °C because it overheated and dried the reaction mixture too much and thus limited the evaporation of the HCl-H 2 O mixture.

3.3. Influence of the Air Flow Rate. To evacuate the HCl-H 2 O mixture out of the reactor, an air circulation on the surface of the reaction mixture was applied. The influence of the air flow rate is presented in Figure 4. The increase of the air flow rate from 60 to 120 L h -1 increased the efficiency of the process. But there was no gain by passing the air flow rate from 120 to 240 L h -1 . So the air flow rate of 120 L h -1 was chosen for the next tests.

Influence of the Water Addition Flow Rate.

The evacuation of HCl is the key parameter for the advancement of the reaction. Since HCl forms with water, an azeotrope containing 20.2 wt % of HCl which has the boiling point at 108.6 °C, the rate of water addition, the temperature of the reaction mixture, and the air flow rate need to be controlled to favor the formation of this azeotrope. Figure 5 shows the influence of the water addition flow rate to the reaction mixture when keeping constant the temperature of the thermal oil bath at 115 °C and the air flow rate at 120 L h -1 , which were the most favorable conditions for the reaction. The increase of the water addition flow rate was not favorable for the reaction under the present experimental conditions. The best results were obtained with the flow rate of 5 g h -1 (or about 0.08 mL min -1 ). This is also the minimum working flow rate of the pump that we used. Another experiment was done, without adding water. The results were similar compared to those of adding 5 g of water per h at the beginning of the reaction. Then, the reaction slowed down because there was no more water in the medium. We consider that the best flow rate of water addition was 5 g h -1 .

3.5. Influence of the Reaction Time and Reproducibility of the Process. Using the reactor N°3, the reaction was not finished after 4 h of reaction. Thus, different tests were carried out for longer reaction times, up to 24 h. The optimal experimental conditions found above were used, including the following: temperature of the oil bath, 115 °C; 500 rpm; air flow rate, 120 L h -1 ; and rate of water addition, 5 or 7.8 g h -1 . The reproducibility of the process was also checked by repeating the reaction several times. Figure 6 shows the results obtained.

For the water addition of 5 g h -1 , the conversion of NaCl to HCl continued up to 24 h of reaction. The conversion profile suggests that the reaction could evolve still. The conversion of 76.5% was obtained after 24 h (test 1, Figure 6a). For the water addition of 7.8 g h -1 , similar results were obtained but the reaction seemed to strongly slow down, taking into account the conversion profile of NaCl to HCl (test 3, Figure 6b). The conversion of NaCl to HCl reached 67.5% after 24 h of reaction. For both the water addition rates, tests of reproducibility allowed confirming that the process was well repeatable (test 2, Figure 6a and tests 4 and 5, Figure 6b).

3.6. Improvement of the Performance of the Synthesis Process. HCl trap allowed channeling and recovering HCl vapor. But vapor condensation took place during the reaction, as illustrated in Figure 7a. In fact, the head of the reactor was not heated, which caused the condensation of the HCl-vapor mixture. This condensation is dramatic because it sent HCl back to the reaction mixture and thus slowed down the reaction.

To the reduction of this condensation, a parametric study was carried out by removing the HCl trap, as illustrated in Figure 7b. In parallel, the consistency of the reaction mixture was also adjusted by manual water addition. In fact, after the first hours of reaction time (0 to 13 h), about 50-70% of orthophosphoric acid were converted into sodium phosphates. The reaction mixture partially lost its initial consistency and could be dried under heating and air flux. This prevented the contact between Cl -anions and protons H + inside the reaction mixture. Manual water addition every hour or every 30 min again allowed for the homogenization of the reaction mixture. After each homogenization, the reaction mixture was progressively dried under heating and air flux before the next manual water addition. By this method, the reaction mixture could reach the composition of HCl-H 2 O azeotrope, which favors the elimination of HCl from the reaction mixture. Note that with the reaction time, the content of chloride in the reaction mixture decreased and the quantity of water required for getting the composition of HCl-H 2 O azeotrope changed. Manual water addition was used for adapting to this evolution. Table 1 summarizes the reaction conditions of this parametric study.

Because HCl was not trapped, we could only analyze the chloride content remaining in the final product, recovered at the end of the reaction (24 h). Figure 8 shows the conversion of NaCl to HCl after 24 h of reaction. The manual water addition (10 g every hour within 8-24 h of reaction) allowed for the increasing performance of the process from 74% to 79% (experiment A vs experiment B). The increase of water addition flow rate by the peristaltic pump from 8 to 15 g h -1 led to the increase of NaCl conversion from 79 to 86% (experiment B vs experiments C and D). When the manual water addition was reduced to 7 instead of 10 g every hour, the performance slightly increased from 86 to 89% (experiment D vs experiment E). Manually adding 5 g of water every 30 min allowed for 90% conversion of NaCl to HCl (experiment F). Finally, the use of NaCl powder having small particles (d 50 of 78 μm) associated with the manual water addition of 10 g every hour led to the conversion of 91% (experiment G). This was also the best performance obtained in this work.

3.7. Analysis of the Final Product. Final product of each test described above was analyzed by ion chromatography and by ICP-AES. At the end of the reaction, the reaction mixture was dissolved with water to form homogeneous aqueous solution. This dissolution took place within less than 1 min under stirring. The resulting solution was quickly analyzed by ion chromatography for determining chloride (Cl -), orthophosphate (PO 4 3-), pyrophosphate (P 2 O 7 4-), and trimetaphos- water additions, see Table 1.

phate (P 3 O 9 3-) species and by ICP-AES for determining sodium and phosphorus. As an example, Table 2 summarizes the analysis results obtained with the final product from experiment G in Figure 8.

Table 2 shows that the quantity of sodium and phosphorus in the final product analyzed by ICP [Na ( f) and P (total) ] reached the initial quantity of sodium and phosphorus introduced in the reactor [PO 43-(i) and Na + (i) , respectively]. So, the mass balance was good for sodium and phosphorus. For chlorine, the mass balance could not be done for this experiment because HCl vapor was not trapped. But for all other experiments using HCl trap, the mass balance was also good (for chlorine). About the distribution of phosphorus-containing products, trimetaphosphate (P 3 O 9 3-) was not observed. In fact, this product is formed at high temperature. [START_REF] Pham Minh | Onestep synthesis of sodium trimetaphosphate (Na 3 P 3 O 9 ) from sodium chloride and orthophosphoric acid[END_REF] Pyrophosphate (P 2 O 7 4-) was only formed at trace amount (2 mmol or 0.6% of the initial quantity of orthophosphoric acid used). Orthophosphate (PO 4 3-) was the principal phosphorus-containing product of the reaction (347 mmol or 99% of the initial quantity of orthophosphoric acid used).

From the results in Table 2, the composition of the final product of the experiment G in Figure 8 was proposed. Taking into account the fact that the initial reaction mixture contained 350 mmol of NaCl and 350 mmol of H 3 PO 4 , this product may contain the following components: (1) 31 mmol of unreacted NaCl (9% of the initial quantity introduced into the reactor).

(2) 31 mmol of unreacted H 3 PO 4 and 316 mmol of NaH 2 PO 4 corresponding to 347 mmol PO 4 3-. We considered that the quantity of unreacted H 3 PO 4 was equal to the quantity of chloride remaining in the reaction mixture (31 mmol). NaH 2 PO 4 must be the main orthophosphate sodium because under acidic pH of the reaction, H 3 PO 4 and H 2 PO 4

-are the main orthophosphate species, and the presence of HPO 4 2-and PO 4

3-must be negligible. 23 (3) 2 mmol of P 2 O 7 2-, which suggest the formation of Na 2 H 2 P 2 O 7 . This product could be formed by the condensation of NaH 2 PO 4 .

Figure 9a sketches the composition of the final product recovered from the experiment G in Figure 8. As previously mentioned, this product was already dissolved in water for ion chromatography and ICP-AES analyses. This solution was later evaporated at room temperature and atmospheric pressure under an extractor hood for the crystallization step. Figure 9b shows the crystals obtained from this evaporation. This product was slightly humid because of the presence of unreacted H 3 PO 4 (9%).

In summary, this work investigated the direct reaction of sodium chloride with orthophosphoric acid. The first objective was to design an alternative to the conventional industrial production of sodium phosphates from orthophosphoric acid and sodium hydroxide or sodium carbonate. From a technical point of view, sodium chloride could react with orthophosphoric acid at moderate temperature, and high conversion and high selectivity in sodium dihydrogen phosphate (NaH 2 PO 4 ) could be obtained. Evaluating the economic viability of this process is further necessary. But it is a complex task which requires multiple criteria and was not the first objective of this study. Thus, we report hereafter some general features about the present process considered, compared to the conventional production way: (1) the reaction of NaOH with H 3 PO 4 is exothermic (Δ f H°: -124.7 kJ mol -1 ), but it is carried out at a temperature around 80-100 °C on an industrial scale. The reaction of NaCl and H 3 PO 4 is also slightly exothermic (Δ f H°: -24.6 kJ mol -1 ) and needs a reaction temperature around 115 °C. So the difference of energy for heating the reaction is small. However, NaOH is industrially produced by electrolysis of NaCl aqueous solution, which is very endothermic (Δ f H°: + 480.6 kJ mol -1 ). So, the energy balance must take into account this electrolysis step, as well as the utility of the final products and byproducts of each synthesis method. The electrolysis of NaCl solution produces H 2 and Cl 2 as byproducts, while the direct reaction of NaCl and H 3 PO 4 produces HCl as the only byproduct. (2) Conventional production of NaH 2 PO 4 from H 3 PO 4 and NaOH must be instantaneous at 80-100 °C, while the reaction of NaCl with H 3 PO 4 needed a long reaction time as observed in the present study. However, as mentioned above, NaOH is produced from NaCl by the specific electrolysis process. The cost and impact of NaOH production must be included into the viability evaluation. (3) The reaction of NaCl and H 3 PO 4 generates HCl as the byproduct, which is corrosive. But technological solution is now available for such a process. Strong acids such as HCl, HNO 3 , and H 2 SO 4 are largely produced on the industrial scale. So, the design of a reactor adapted to the presence of HCl must be possible.

CONCLUSIONS

For the first time, we were successful in the synthesis of sodium orthophosphate from the reaction of NaCl with orthophosphoric acid under moderate conditions. The reaction was studied in the aqueous medium, without organic solvent. The evacuation of HCl was found to be crucial for the reaction. This evacuation could be controlled by different parameters including: reactor geometry, reaction temperature, air flow rate, water addition, and reaction time. The highest conversion of NaCl to HCl obtained in this work reached 91% after 24 h of reaction. Further work will be concentrated on the optimization of the synthesis process to get higher elimination yield of chloride, as well as on the competition of the present process versus standard neutralization of H 3 PO 4 by NaOH or Na 2 CO 3 .

Figure 1 .

 1 Figure 1. Illustration of three glass reactors used in this work.

Figure 2 .

 2 Figure 2. Comparison of three glass reactors. Reaction conditions: 500 rpm; 80 °C; air flow rate, 120 L h -1 ; and water addition, 5 g h -1 .

Figure 3 .

 3 Figure 3. Influence of reaction temperature on the conversion of NaCl to HCl. Reaction conditions: air flow rate, 120 L h -1 ; 500 rpm; and flow rate of water addition, 5 g h -1 .

Figure 4 .

 4 Figure 4. Influence of the air flow rate on the conversion of NaCl to HCl. Reaction conditions: temperature of the oil bath, 115 °C; 500 rpm; and flow rate of water addition, 5 g h -1 .

Figure 5 .

 5 Figure 5. Influence of the flow rate of water addition on the conversion of NaCl to HCl. Reaction conditions: temperature of the oil bath, 115 °C; 500 rpm; and air flow rate, 120 L h -1 .

Figure 6 .

 6 Figure 6. Influence of the reaction time on the conversion of NaCl to HCl. Reaction conditions: temperature of the oil bath, 115 °C; 500 rpm; air flow rate, 120 L h -1 ; and flow rate of water addition (a) 5 g h -1 or (b) 7.8 g h -1 .

Figure 7 .

 7 Figure 7. (a) Illustration of HCl condensation during the synthesis reaction with HCl trap and (b) reactor without HCl trap.

Figure 8 .

 8 Figure 8. One-step synthesis of NaH 2 PO 4 from NaCl and H 3 PO 4 at atmospheric pressure and without the HCl trap. Reaction conditions: 500 rpm; 24 h of reaction; Thermal oil's temperature, 145 °C ; and water additions, see Table 1.

Figure 9 .

 9 Figure 9. (a) Composition (mmol) of the final product obtained by the experiment G in Figure 8 and (b) illustration of the solid obtained after crystallization of this product.

Table 1 .

 1 Reaction Conditions for the Synthesis of NaH 2 PO 4 from NaCl and H 3 PO 4 using Reactor N°3 without HCl Trap; the Mean Diameter d 50 of NaCl was 375 μm for All Experiences, except for the Experience G

	experiment	continuous addition of water by the peristaltic pump (g h -1 )	manual water addition
	A	8	without manual water addition
	B	8	10 g every hour between 8 and 24 h of reaction
	C	15	10 g every hour between 8 and 24 h of reaction
	D	15	10 g every hour between 13 and 24 h of reaction
	E	15	7 g every hour between 13 and 24 h of reaction
	F	15	5 g every 30 min between 13 and 24 h of reaction
	G	15	similar reaction conditions to D but with NaCl of d 50 of 78 μm instead of 375 μm

Table 2 .

 2 Analysis of Final Product of Experiment G in Figure 8 by Ion Chromatography and by ICP-AES

						amount
	product	species or element analyzed			(mmol)
	composition of the initial reaction mixture	orthophosphate initially introduced into the	PO 4	3-(i)	350
		reactor
		chloride initially introduced into the reactor	Cl -(i)	350
		sodium initially introduced into the reactor	Na +	(i)	350
	analysis of the final product by ion chromatography	chloride remaining in the reactor	Cl -( f)	32
		orthophosphate remaining in the reactor	PO 4	3-( f)	347
		pyrophosphate formed in the reactor	P 2 O 7 4-( f)	2
		tripolyphosphate remaining in the reactor	P 3 O 9 3-( f)	0
	analysis of the final product by ICP-AES	sodium remaining in the reactor	Na (f)	349
		total phosphorus in the reactor	P (total)	351

■ AUTHOR INFORMATION

Corresponding Author *E-mail: doan.phamminh@mines-albi.fr. Tel: +33563493258. Fax: +33563493243.

Notes

The authors declare no competing financial interest.

■ ACKNOWLEDGMENTS

The authors gratefully acknowledge colleagues at the RAPSODEE centre for technical help.


Augustina Ephraim
email: augustina.ephraim@mines-albi.fr.

Victor Pozzobon

Olivier Louisnard

Doan Pham

Ange Nzihou

Patrick Sharrock

Simulation of biomass char gasification in a downdraft reactor for syngas production

Keywords: biomass waste, gasification, syngas, model simulation, downdraft reactor, OpenFOAM

ou non, émanant des établissements d'enseignement et de recherche français ou étrangers, des laboratoires publics ou privés.

Introduction

As global energy demand is projected to rise by nearly 56% to 865 EJ between 2010 and 2040, syngas will become increasingly important for electric power generation, process heat and liquid fuel production. [START_REF]International Energy Oulook[END_REF] Syngas is a mixture of hydrogen (H 2 ) and carbon monoxide (CO) produced from the gasification of carbonaceous feedstock. As biomass is the only renewable carbon-based energy source, syngas produced from waste biomass (biosyngas) will be a key-intermediate in the future production of renewable fuels to meet national renewable energy and global carbon-dioxide-emission reduction targets. [START_REF] Van Der Drift | Bio-syngas: key intermediate for large scale production of green fuels and chemicals[END_REF] Examples of waste biomass materials include residues from forest and wood processing industries such as wood chips, saw dust and bark, and agricultural residues such as straw, bagasse, and husk. [START_REF] Alakangas | CEN technical specification for solid biofuels-fuel specification and classes[END_REF] Pyrolysis is the thermal conversion of biomass under inert or oxygen-deficient atmosphere at temperatures between 500 and 8008C. The three primary products of pyrolysis are biooil, char and gas. At high temperatures (>8008C), the char produced from pyrolysis can be gasified with gasification agents such as steam (H 2 O), carbon-dioxide (CO 2 ), and oxygen (O 2 ) in air to produce syngas. The most important industrial applications of syngas, so far practiced commercially or under research studies, can be summarized as follows [START_REF] Boerrigter | Syngas production and utilisation[END_REF][START_REF] Basu | Biomass Gasification and Pyrolysis: Practical Design and Theory[END_REF] :

1. Electricity production in steam cycles, gas engines, turbines (combined cycle), and fuel cells, 2. Chemical synthesis of hydrogen (for refineries), Synthetic Natural Gas and ammonia (mainly for fertilizer production), 3. Transport fuel synthesis from Gas-to-Liquids processes including Fischer-Tropsch (FT) diesel and methanol/ dimethyl ether.

The choice of syngas end-use mainly depends on its composition (H 2 /CO ratio) as shown in Table 1. [START_REF] Butterman | CO2 as a carbon neutral fuel source via enhanced biomass gasification[END_REF] Several operating parameters influence the composition of syngas, the most important of which are the gasification agent, reactor temperature, biomass properties and the use of catalysts in the gasification process. [START_REF] Arena | Process and technological aspects of municipal solid waste gasification. A review[END_REF][START_REF] Belgiorno | Energy from gasification of solid wastes[END_REF][START_REF] Devi | A review of the primary measured for tar elimination in biomass gasification processes[END_REF] As a result of the highly heterogeneous nature of waste biomass feedstock, the prediction of syngas H 2 /CO ratio from the physical and chemical properties of biomass char has proven to be difficult. [START_REF] Klaas | Optimisation of hydrogen production by steam reforming of chars derived from lumber and agricultural residues[END_REF] Consequently, recent research works have focused on studying the influence of temperature, gasification agents and catalysts on syngas composition to manipulate the syngas ratio and better control the gasification process. [START_REF] Butterman | CO2 as a carbon neutral fuel source via enhanced biomass gasification[END_REF][START_REF] Klaas | Optimisation of hydrogen production by steam reforming of chars derived from lumber and agricultural residues[END_REF][START_REF] Yang | Utilization of CO2 and biomass char derived from pyrolysis of Dunaliella salina: the effects of steam and catalyst on CO and H2 gas production[END_REF][START_REF] Yang | Effect of model and operating parameters on air gasification of char[END_REF][START_REF] Yan | Hydrogen-rich gas production by steam gasification of char from biomass fast pyrolysis in a fixed-bed reactor: influence of temperature and steam on hydrogen yield and syngas composition[END_REF][START_REF] Chaudhari | Steam Gasification of biomass-derived char for the production of carbon monoxide-rich synthesis gas[END_REF][START_REF] Moazami | Modelling of a fixed bed reactor for Fischer-Tropsch synthesis of simulated N2-rich syngas over Co/SiO2: hydrocarbon production[END_REF] This ensures that the biomass feedstock has a lesser impact on the syngas produced and thus offers greater fuel flexibility to meet the requirements of the desired syngas application. [START_REF] Butterman | CO2 as a carbon neutral fuel source via enhanced biomass gasification[END_REF] In view of engineering design and process optimization of a biomass gasifier for syngas production, an extensive investigation of the plant behavior depending on various operating parameters is required. [START_REF] Hamel | Mathematical modelling and simulation of bubbling fluidised bed gasifiers[END_REF] Performing experiments at large scales are often problematic mainly due to their associated safety risks and high costs. Hence, developing a mathematical model that gives a good representation of the complex chemical and physical phenomena occurring in the gasifier is desirable as it allows fewer and better experiments to be performed with minimal temporal and financial costs. In addition, a simulation model enables prediction of the evolution of these complex phenomena which is essential for optimizing the reactor design and its operating conditions. [START_REF] Ahmed | Mathematical and computational approaches for design of biomass gasification for hydrogen production: a review[END_REF] For models of biomass char gasification, the vast majority of those presently found in literature are kinetic models obtained from the study of char reactivity in different atmospheres and temperature ranges, [START_REF] Koenig | Evidence for two-site model of char gasification by carbon dioxide[END_REF][START_REF] Guizani | Influence of temperature and particle size on the single and mixed atmosphere gasification of biomass char with H2O and CO2[END_REF][START_REF] Fermoso | Kinetic models comparison for non-isothermal steam gasification of coal-biomass blend chars[END_REF][START_REF] Kramb | Modeling biomass char gasification kinetics for improving prediction of carbon conversion in a fluidized bed gasifier[END_REF][START_REF] Septien | Steam gasification of char from wood chips fast pyrolysis: development of a semiempirical model for a fluidized bed reactor application[END_REF] disregarding diffusive and convective heat and mass transfer, and thus neglecting temperature and composition gradients in the sample. Conversely, computational fluid dynamics (CFD) modeling techniques applied to biomass char gasification are scarce and are currently restricted to either coal gasification or biomass combustion/gasification in entrained flow reactors. [START_REF] Ahmed | Mathematical and computational approaches for design of biomass gasification for hydrogen production: a review[END_REF] CFD programs allow the coupling of fluid flow, heat and mass transfer to chemical reactions and other related phenomena, by numerically solving the set of governing conservation equations: mass, heat, and momentum. [START_REF] Wang | CFD studies on biomass thermochemical conversion[END_REF] To the best of our knowledge, the most recent experimentally validated CFD model in literature on char gasification is that of Teixeira et al. [START_REF] Teixeira | Gasification of continuous wood char bed: modelling and experimental approach[END_REF] who developed a steady state, one-dimensional (1-D) model of wood char gasification in a downdraft reactor.

In this article, modifications are made to the latter CFD model in an attempt to improve its prediction of syngas composition. An original in-house built solver based on Open-FOAM is used to simulate the model, and the results are validated against the experimental data of Teixeira et al. [START_REF] Teixeira | Gasification of continuous wood char bed: modelling and experimental approach[END_REF] OpenFOAM is an open source CFD software package which offers users complete freedom to customize and extend its existing functionality. OpenFOAM's capacity to solve complex CFD problems is comparable to that of reputed powerful close sourced software such as ANSYS Fluent. Although OpenFOAM has a large user community across most areas of engineering and science, it has been scarcely used for simulating biomass gasification. The only biomass gasification solver found in literature is biomassGasificationFOAM which was developed by Kwiatkowski et al. [START_REF] Kwiatkowski | Pyrolysis and gasification of single biomass particle-new openFoam solver[END_REF] to simulate the pyrolysis and gasification of a single biomass particle. Thus until now, no OpenFOAM solver has been developed to simulate the gasification of biomass at the reactor scale in a downdraft fixed bed gasifier. Next in this article, a sensitivity analysis is performed to study the influence of temperature and a mixed atmosphere (O 2 , H 2 O, CO) on char conversion, bed temperature profile and H 2 and CO production (H 2 /CO ratio). This sensitivity analysis will enable clarification of the complex reaction mechanisms involved in mixed atmosphere gasification, and an evaluation of the most suitable operating parameter(s) for optimizing syngas composition, to meet the requirements of the desired industrial application.

Modeling Approach

A schematic view of a downdraft reactor is shown in Figure 1a. The biomass particles are fed at the top of the reactor and slowly flow to the bottom where the residual ash is withdrawn. The gasification agents are injected through the sides of the reactor. During their downward flow, the biomass particles undergo the following main processes: drying, pyrolysis, combustion and reduction. Finally, the produced gas is withdrawn from the bottom of the reactor.

Our work focuses on modeling the gasification of char produced from the pyrolysis of wood biomass by considering reactions of char with the gasification agents, air (O 2 1 N 2 ), steam (H 2 O) and carbon dioxide (CO 2 ). The reactions considered in this model and their associated heats of reaction at 258C are

C 1 O 2 ! CO 2 Combustion 2394 MJ=kmol (1) C 1 CO 2 ! 2CO Boudouard 1172 MJ=kmol (2) C 1 H 2 O ! CO 1 H 2 Steam gasification1131 MJ=kmol (3) CO 1 H 2 O $ CO 2 1 H 2 Water-gas shift WGS ð Þ41 MJ=kmol ( 4
)
where the N 2 in air has been assumed to remain inert. Figure 1b shows the char bed in the gasification zone of the reactor in which the equations of mass, heat and momentum are solved. The boundary conditions for these equations are applied at the bed inlet and outlet (z 5 0 and z 5 H, respectively).

Assumptions

The following assumptions were made to model the gasification of wood char in the downdraft reactor:

1. Steady-state operation as most industrial downdraft gasifiers operate for long periods with few interruptions and the feed and withdrawal rates of the solid and gas phases have relatively low fluctuations.

2. Laminar flow regime for the gas phase due to the low velocities usually encountered in fixed bed reactors.

3. Ideal gas mixture as most downdraft gasifiers operate at atmospheric pressure.

4. Char bed is a homogeneous porous media so that all variables are assumed continuous in space and time.

5. At the pore scale, there is no heat transfer resistance between the solid and the gas phases. This assumption is based on the experimental findings of Van de Steene et al. [START_REF] Van De Steene | Gasification of woodchip particles: experimental and numerical study of char-H2O, char-CO2, and char-O2 reactions[END_REF] [Color figure can be viewed in the online issue, which is available at wileyonlinelibrary.com.]

The complex medium can therefore be represented with a single temperature field (T g 5 T c 5 T). 6. Heat transfer by radiation is neglected due to the relatively large flows of gas and solid which renders convective and conductive heat transfer predominant.

7. Constant bed porosity, tortuosity and particle diameter are assumed along the char bed as the evolution of these parameters with char conversion is very difficult to determine.

Mathematical equations and boundary conditions

Continuity Equation for Gas Phase Coupled to Momentum Balance (Darcy's Law). The continuity equation for the gas phase is written as

r Á ðU g C g Þ5 X 6 j51 r j (5)
where U g is the superficial gas velocity, C g is the total gas concentration and r j is the mass source of gas species j, where j belongs to CO, CO 2 , H 2 O, O 2 , H 2 . By considering the gas as an ideal mixture, C g can be expressed as a function of bed pressure P, and temperature T, by

C g 5 P RT ( 6
)
where R is the universal gas constant.

As with several other fixed bed models, [START_REF] Lamarche | Modelling of an indirectly heated fixed bed pyrolysis reactor of wood: transition from batch to continuous staged gasification[END_REF][START_REF] Anca-Couce | How to determine consistent biomass pyrolysis kinetics in a parallel reaction scheme[END_REF][START_REF] Vîjeu ; R, Gerun | Dimensional modelling of wood pyrolysis using a nodal approach[END_REF] Darcy's law is used to describe the average flow of gas through the porous char bed as

U g 52 K l rP ( 7
)
where K is the bed permeability and l is the gas viscosity. Darcy's law can be coupled to the continuity equation by substituting Eqs. 6 and 7 into Eq. 5 to finally give

r Á 2 P RT K l rP 5 X 6 j51 r j (8)
Mass Balance Equation for Each Gas Phase Species. The mass balance for each chemical species j in the gas phase is given by r Á ðU g C g;j Þ5r Á ðD Ã j;N2 rC g;j Þ1r j

The left-hand side of Eq. 9 represents the contribution of the convective mass transfer and the first term on the righthand side represents diffusive transfers. Here we have assumed that since N 2 is the dominant gas in the gas phase, the effective diffusion coefficient of species j in the gas phase is equivalent to the effective diffusion coefficient D Ã j;N2 of species j in pure N 2 . This effective diffusion coefficient is a function of the bed porosity e, and the bed tortuosity s 30 and the bulk diffusion coefficient between species j and N 2

D Ã j;N2 5 e s D j;N2 (10)
The dependence of the bulk diffusion coefficient D j;N2 on temperature and pressure is expressed according to the Chapaman-Enskog formula

D j;N2 5D j;N2 T ref ; P ref ð Þ P ref P T T ref 3=2 ( 11
)
where

D j;N2 T ref ; P ref ð
Þis the diffusivity of species j in nitrogen (N 2 ) gas solvent under reference conditions. Table 2 lists the expression for the diffusion coefficient for each gaseous species j.

The bed tortuosity is approximated by 32 s5120:41lnðeÞ

Finally, the last term r j in the RHS of Eq. 9 is the production rate of species j, and will be detailed hereafter.

Mass Balance Equation for the Solid Phase (Char). The mass balance equation for the solid phase is given by

r Á ðU c C c Þ5r c ( 13
)
where U c and C c are the velocity and concentration of char, respectively, and r c is the sink term for the consumption of char by the chemical reactions.

The velocity of char in a downdraft reactor can decrease significantly as a result of bed compaction. Bed compaction is a consequence of the reduction of particle size due to carbon conversion and various mechanical phenomena such as fragmentation and particle rearrangement. [START_REF] Teixeira | Gasification of char from wood pellets and from wood chips: textural properties and thermochemical conversion along a continuous fixed bed[END_REF] Equation 13 can be coupled to the char bed compaction equation which expresses the solid phase velocity as a function of char conversion

X U c X ð Þ5U c;0 f X ð Þ (14)
where f(X) is an empirical function developed by Teixeira et al. [START_REF] Teixeira | Gasification of continuous wood char bed: modelling and experimental approach[END_REF] and determined for wood char chips f X ð Þ521:03 Á 10 24 X 2 14:25

Á 10 24 X11 (15)
The conversion of char is calculated from the molar flow rates of the gaseous carbon species produced (CO and CO 2 )

X5 _ n c j z50 2 _ n c _ n c j z50 5 _ n CO 1 _ n CO2 2 _ n CO2 j z50 2 _ n CO j z50 _ n c j z50 (16)
Reaction Kinetics. The reaction rates for the heterogeneous reactions m i (Eqs. 1-3) are determined by

v i 5x i C c;0 ( 17
)
where C c,0 is the initial concentration of char and the notation x i denotes the char conversion rate for each heterogeneous reaction (1)-( 3). These char conversion rates have been correlated by Teixeira and coworkers to the experimental parameters, temperature T and partial pressure of the reactant gas P j

The functions of the char particle conversion rates in the three reactant gases are given in Eqs. 19

The properties of wood char used in this model are e p 55:5 3 10 23 m; e p 50:75 and the pre-exponential factors are A ðO2Þ 5 1:1 Á 10 9 s 21 atm 20:6 ; A CO2 51:2 Á 10 8 s 21 atm 20:7 ; and A H2O 5 3:55 Á 10 3 s 21 atm 20:8 , respectively.

For the water-gas shift reaction (WGS), we used a simple reversible rate expression for CO conversion which was proposed by Moe 35 v 4 5k WGS P CO P H2O 12 P CO2 P H2 K eq WGS P CO P H2O M CO (22) k WGS 51:85310 [START_REF] Kwiatkowski | Pyrolysis and gasification of single biomass particle-new openFoam solver[END_REF]

where k WGS (mol g 21 min 21 ) and K eq WGS are the rate constant and equilibrium rate constant respectively.

From the knowledge of the reaction rates v i , i 5 1,. . .4, the production term r j of chemical species j appearing in Eqs. 9 and 13 can be expressed by

r j 5 X 4 i51 m ij v i (25)
where v ij denotes the (algebraic, positive for product, negative for reactants) stoichiometric coefficient of species j in reaction i. Table 4 details the expression for the production terms of the chemical species. Energy Balance Equation. With the assumption of local thermal equilibrium between the gas and solid phases, the equation of energy conservation writes

U g C g c p;g M g 1U c C c c p;c M c À Á rT5r Á ðk b rTÞ1 _ Q r 1 _ Q loss ( 26
)
where convective heat transfer is represented by the term on the left-hand side of Eq. 26 and conductive heat transfer by the first term of the right-hand side. The effective thermal conductivity k b in the porous media is evaluated by assuming that heat conduction in the solid and gas phase takes place in parallel which follows from the fifth assumption that no heat transfer occurs between the solid and gas phases 36

k b 5ð12eÞk c 1ek g ( 27
)
Table 5 details the expression of the thermal conductivities of the chemical species in the two phases as a function of temperature.

In Equation 26, _ Q is the heat source term which can be expressed in function of the heats DH r;i of reactions ( 1)-( 4)

Q r 52 X 4 i51 v i DH r;i (28)
Conversely, the heat lost to the surrounding by convection from the outer wall of the reactor Q loss , was estimated to be 37.7 kWm 23 by Teixeira [START_REF] Teixeira | Gaz eification de charbon de granules de bois: comportement thermochimique et m ecanique d'un lit fixe continu[END_REF] which is determined from Newton's law of cooling as

Q loss 5 hAðT w 2T 1 Þ V ( 29
)
where A and V denote the outer surface area and volume of the reactor, respectively. The average wall temperature T w and ambient temperature T 1 were experimentally found to be 628C and 208C, respectively, and the heat transfer coefficient h is 10 Wm 22 K 21 . Boundary Conditions. The system of differential equations formed by Eqs. 5-26 can be solved with the following set of boundary conditions that specify concentration, temperature and pressure values or their derivatives at the inlet (z 5 0) and outlet (z 5 H) of the char bed:

Concentration of species in gas and solid phases Inlet C j j z50 5C j;0 (30)

r C C c;0 2x 1 2x 2 2x 3 ½ r CO2 C c;0 x 1 2x 2 ½ 1v 4 r H2O 2C c;0 x 3 2v 4 r O2 2C c;0 x 1 r CO C c;0 2x 2 1x 3 ½ 2v 4 r H2 C c;0 x 3 1v 4 Outlet n Á rC j j z5H 50 (31)
Bed temperature

Inlet Tj z50 5T 0 (32) Outlet n Á rTj z5H 50

Bed pressure

Inlet 2 K l Á rPj z50 5U g;0 (34)
Outlet Pj z5H 5P atm (35)

OpenFOAM simulation

The system of partial differential equations discussed in the previous section was solved using the finite volume discretization method employed by OpenFOAM. The chemical and thermal properties data were obtained from NIST-JANAF Thermochemical Tables 39 andTeixeira. [START_REF] Teixeira | Gaz eification de charbon de granules de bois: comportement thermochimique et m ecanique d'un lit fixe continu[END_REF] A mesh convergence study was performed and an optimum value of 325 cells was found.

Results and Discussion

Validation of simulation model

The model predictions have been validated against three different sets of experimental results: two different feedstocks with the same operating conditions (Table 6), and two different experimental conditions with the same feedstock (Table 7). The results are presented in Table 8. It can be seen that the model predictions are in good agreement with all three experimental observations. Discrepancies remain regarding CO and H 2 concentrations at the outlet of the bed; nevertheless the model captures the trend well.

A more detailed evaluation of the model's ability to predict the evolution of the bed behavior along the bed height was made by comparing the predicted results with the experimental ones for char chip gasification provided by Teixeira et al. [START_REF] Teixeira | Gasification of continuous wood char bed: modelling and experimental approach[END_REF] The results are displayed in Figure 2 which reveal a highly reactive zone at the top of the bed (z < 10 cm) where the bed temperature decreases sharply to 8758C due to the predominant endothermic char gasification reactions involving H 2 O and CO 2 . Below this zone, the low temperatures slow down the char gasification reactions which in turn decrease the production of H 2 and CO. To estimate the accuracy of our simulation results (w is ) with respect to the experimental data (w i e), the sum squared method was used Mean errorð%Þ5100

ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ffi X N i51 wie2wis wie 2 N v u u t (36)
The error analysis on our simulation results reveals a discrepancy of 4% for char conversion, 2% for bed temperature, 11% for H 2 and 4% for CO. These discrepancies are satisfactory given that the estimated experimental error reported by Teixeira et al. [START_REF] Teixeira | Gasification of continuous wood char bed: modelling and experimental approach[END_REF] is 11%.

Figure 2 also shows a discrepancy between our simulation results and those of Teixeira, especially on the concentration fields of H 2 and CO (Figures 2c,d). It appears that the improved prediction of CO concentration by our model may be linked to our choice of WGS kinetics (Eqs. 22-24), which differs from the one adopted by Teixeira and coworkers. The WGS kinetics used by Teixeira et al. [START_REF] Vassilev | An overview of the behaviour of biomass during combustion: part I. Phase-mineral transformations of organic and inorganic matter[END_REF] was for an uncatalyzed reaction (Eq. 37 and Table 9) Value not provided by the author and hence is estimated to be the same as that for char chips used by Teixeira.

v 4 5 k WGS RT ð Þ 2 P CO P H2O 2 P CO2 P H2 K eq WGS ( 37
)
where k WGS and K eq WGS follow Arrhenius laws (Table 9). On the contrary, our model uses Moe's kinetics [START_REF] Moe | Design of water-gas shift reactors[END_REF] which accounts for WGS catalysis by iron oxide (Fe 2 O 3 ) in the reactor bed. This choice is based on the knowledge that wood char contains ash which is composed of up to 9.5 wt % Fe 2 O 3 [START_REF] Smith | A review of the water gas shift reaction kinetics[END_REF] and that commercially available high temperature WGS catalysts are typically manufactured from Fe 2 O 3. [START_REF] Van De Steene | A new experimental continuous fixed bed reactor to characterise wood char gasification[END_REF] In the reactor entrance zone of 10 cm height where 80% of the char is converted to gas (Figure 2a), the remaining solid is essentially ash. Thus we expect that in this zone, the ash containing Fe 2 O 3 will catalyze the WGS reaction and thus increase CO conversion. Although our use of Moe's WGS kinetics [START_REF] Moe | Design of water-gas shift reactors[END_REF] significantly improved CO prediction (Figure 2d), the opposite held for H 2 (Figure 2c). This discrepancy for H 2 can be explained from the observation that in Figures 2c, and2d, close to the reactor exit, H 2 is still being produced while CO is consumed. This is because the WGS reaction rate does not reach equilibrium as shown in Figure 3 by the positive value of the reaction rate at the reactor exit. Hence, given a longer reactor, we expect that the error observed for H 2 will diminish as the WGS reaction nears equilibrium. Nonetheless, it is clear that the choice of the WGS kinetics plays a key role in predicting the final syngas composition.

Sensitivity analysis

The aim of this analysis is to investigate the influence of the bed inlet temperature on the one hand, and of the gas composition on the other hand, on the bed temperature profile T, char In all graphs, the experimental data provided by Teixeira et al. [START_REF] Teixeira | Gasification of continuous wood char bed: modelling and experimental approach[END_REF] are represented by triangles, and the predicted values from OPENFOAM and COMSOL [START_REF] Teixeira | Gasification of continuous wood char bed: modelling and experimental approach[END_REF] simulations are represented by solid and dashed lines, respectively. [Color figure can be viewed in the online issue, which is available at wileyonlinelibrary.com.] conversion X and syngas composition (CO, H 2 ). Char from maritime pine wood chips is the chosen feedstock whose composition and properties are provided by Teixeira et al. [START_REF] Teixeira | Gasification of continuous wood char bed: modelling and experimental approach[END_REF][START_REF] Teixeira | Gasification of char from wood pellets and from wood chips: textural properties and thermochemical conversion along a continuous fixed bed[END_REF] and are shown in Table 6. The operating conditions used for this study are displayed in Table 10.

To perform the sensitivity analysis, the following operating parameters were varied:

Inlet bed temperature, T 0 : 900-1000 8C O 2 concentration, C O2;0 : 0-3 vol % Steam concentration, C H2O;0 : 5-20 vol % CO 2 concentration, C CO2 ;0 : 5-20 vol % For each set of operating parameters, their influence is investigated by analyzing the char conversion, bed temperature profile and the production of H 2 and CO along the char bed. Moreover the evolution of the different reaction rates is examined to clarify the mechanism involved.

Effect of Inlet Bed Temperature. Figure 4 shows the simulation results for char conversion, temperature profile, H 2 and CO concentration for different inlet bed temperatures. These results show that an increase in the bed inlet temperature from 900 to 1000 8C increases overall char conversion and CO concentration by 27 % and 33 % respectively. It also increases to a lesser extent the concentration of H 2 and outlet temperature (17% and 148C, respectively). Consequently, the H 2 /CO ratio decreases from 0.7 to 0.6. These results can be explained by observing the influence of inlet bed temperature on the reaction rates as shown in Figure 5. Figures 5a and5b, show that in the 5 cm entrance zone of the reactor, the increase in bed temperature causes a drastic rise of the endothermic steam gasification and Boudouard reactions (Eqs. 2 and 3). This leads to a strong increase in CO production and an insignificant rise in H 2 production. Higher temperatures in this reaction zone also favor the production of CO to the detriment of H 2 by shifting the WGS equilibrium position to the left (Eq. 4). This effect of inlet bed temperature on CO enrichment in syngas is consistent with the observations made by Yang et al. [START_REF] Yang | Utilization of CO2 and biomass char derived from pyrolysis of Dunaliella salina: the effects of steam and catalyst on CO and H2 gas production[END_REF] from char gasification experiments at lower temperatures (500-800 8C). Producing CO-rich syngas (low H 2 /CO ratio) is desirable for FT synthesis. However this may not justify the indirect heating of the reactor to temperatures above 900 8C due to the associated high energy costs and the relatively small gain in lowering in the H 2 /CO ratio.

Effect of O 2 Concentration. The simulation results for char conversion, temperature profile, H 2 and CO concentration for different O 2 concentration of the inlet gas stream are shown in Figure 6. The results show that an increase in O 2 concentration from 0 to 3 vol % increases the overall char conversion by 48%, the bed outlet temperature by 86 8C and the production of H 2 and CO by 20 % and 133 % respectively. Thus the H 2 /CO ratio falls from 0.8 to 0.4. To understand this phenomenon, it is useful to examine the evolution of the different reaction rates (Figure 7). In an entrance zone of 10 cm height, the increase in O 2 concentration strongly increases the combustion rate which produces more CO 2 and raises the bed temperature (Eq. 1). The production of CO 2 favors in turn the Boudouard reaction and thus more CO is produced (Eq. 2). Next, the combination of CO 2 and bed temperature increase shifts the WGS equilibrium to the left which increases production of H 2 O (Eq. 4) and further increases production of CO. Finally, the steam gasification reaction (Eq. 3) is favored by the increase of H 2 O and thus also contributes to the production of CO and H 2 . Hence the simulation results show that a relatively small increase in O 2 concentration (0-3 vol %) strongly increases char conversion to favor CO production and also increases bed temperatures. Yang et al. [START_REF] Yang | Effect of model and operating parameters on air gasification of char[END_REF] have reported a similar effect of O 2 in air on CO production and bed temperature when air velocity is increased during char gasification. Thus our results show that raising the O 2 concentration is beneficial to provide direct heating to the reactor and to produce syngas with a low H 2 /CO ratio for FT synthesis. Of course, a trade-off must be found between this benefit of increasing O 2 and the additional costs incurred from operating at higher air flow rates or producing oxygen-enriched air from air separating units.

Effect of H 2 O Concentration. Figure 8 shows the numerical results for char conversion, temperature profile, H 2 and CO concentration for different H 2 O concentrations of the inlet gas stream. It can be seen that the increase in H 2 O concentration from 5 to 20 vol % raises the overall char conversion by 17 % and the H 2 production by 150 %, whereas it decreases the CO production by 33 % and the bed outlet temperature by 6 8C. As a result, the H 2 /CO ratio increases from 0.3 to 1.3. Figure 9 shows that in an entrance zone of 6 cm height, the increase in H 2 O concentration boosts the steam gasification reaction (Eq. 3), thereby producing more CO and H 2 . Much of this CO produced is consumed by the WGS reaction to favor H 2 production (Eq. 4). As steam gasification is an endothermic reaction, an increase in its reaction rate decreases the bed temperature 6.

[Color figure can be viewed in the online issue, which is available at wileyonlinelibrary.com.] which in turn decreases the rate of the Boudouard reaction and the production of CO. Thus an increase in H 2 O concentration in the gas at the bed inlet results in a significant increase in H 2 production to the detriment of CO production. Steam gasification experiments of biomass char in a fixed bed reactor by Yan et al. [START_REF] Yan | Hydrogen-rich gas production by steam gasification of char from biomass fast pyrolysis in a fixed-bed reactor: influence of temperature and steam on hydrogen yield and syngas composition[END_REF] have revealed that increasing steam flow rate from 0 to 0.165 g/min/g of biomass char increased H 2 /CO ratio from 0.88 to 3.74. Consequently, syngas produced from steam gasi-fication can be used in gas turbine combustion and Solid oxide fuel cells (SOFC). Effect of CO 2 Concentration. Figure 10 shows the numerical results for char conversion, temperature profile, H 2 and CO concentration for different CO 2 concentrations of the inlet gas stream. The results reveal that the increase in CO 2 concentration from 5 to 20 vol % increases the overall char conversion by 74 %, the outlet bed temperature by 27 8C and the CO [Color figure can be viewed in the online issue, which is available at wileyonlinelibrary.com.] production by 10 %, whereas it decreases the H 2 production by 35 %. As a result, the H 2 /CO ratio decreases from 0.7 to 0.4. Figure 11 shows that between the bed inlet and a height of 10 cm, the increase in CO 2 concentration from 5 to 10 vol % increases the Boudouard reaction rate and shifts the WGS equilibrium position to the left, thus it increases CO (8 %) and decreases H 2 (12 %). As both of these reactions are endother-mic, the bed temperature decreases. However when the initial CO 2 is 20 vol %, the decrease in bed temperature is significant enough to decelerate the steam gasification reaction between the bed height of 10 cm and 20 cm from the inlet, as shown in Figure 11a. This decrease in the steam gasification rate not only decreases the CO and H 2 production as predicted by Eq. 3, but also increases the bed temperature as observed in [Color figure can be viewed in the online issue, which is available at wileyonlinelibrary.com.] [Color figure can be viewed in the online issue, which is available at wileyonlinelibrary.com.] [Color figure can be viewed in the online issue, which is available at wileyonlinelibrary.com.] [Color figure can be viewed in the online issue, which is available at wileyonlinelibrary.com.]

Conclusions

In this article, a 1-D CFD simulation model developed using OpenFOAM was presented to simulate the gasification of wood char in a downdraft reactor. Good quantitative agreement between the numerical results of the simulation and published experimental data was achieved. A sensitivity analysis was also conducted to investigate the influence of the bed inlet temperature and the inlet gas composition on char conversion, temperature profile and concentrations of CO and H 2 in the product gas. A summary of the results obtained from the sensitivity analysis is shown in Table 11. From these results, the following conclusions can be drawn:

A rise in inlet bed temperatures above 900 8C has a relatively small influence on char conversion and syngas H Although increasing the inlet CO 2 concentration from 5 to 10 vol % increases the char conversion and CO production, further increasing the inlet CO 2 concentration from 10 to 20 vol % actually lowers the syngas quality because the net increase of CO is small and H 2 production is further decreased. This restricts the recycling of industrial CO 2 as an agent for controlling syngas composition. The in-house built OpenFOAM solver used in this work is the first of its kind to simulate the gasification of biomass at the reactor scale in a downdraft gasifier. The development of this simulation model which is capable of predicting the composition of syngas with varying operating conditions provides essential knowledge for optimizing the design and operation of downdraft gasifiers.

Future Work

The model presented in this article can be modified to simulate the gasification of raw biomass in a stratified downdraft reactor. To do so, submodels describing the drying and pyrolysis stages of the gasification processes could be incorporated. Furthermore, as a result of the low computation time for each simulation run (30 s average), optimization techniques can be used to determine optimal reactor design and operating conditions with respect to the syngas H 2 /CO ratio desired.

Figure 1 .

 1 Figure 1. (a) Schematic view of a downdraft reactor and (b) Model representation of the char gasification zone.

Figure 2 .

 2 Figure 2. Experimental and numerical results for char conversion (a), temperature profile (b), H 2 concentration (c) and CO concentration (d).

Figure 3 .

 3 Figure 3. Evolution of WGS reaction rate along the char bed at the operating conditions specified in Table6.

Figure 4 .

 4 Figure 4. Influence of inlet temperature (9008C-solid line, 9508C-dashed line, 10008C-dash-dotted line) on char conversion (a), temperature profile (b), H 2 concentration (c), and CO concentration (d).

Figure 5 .

 5 Figure 5. Influence of inlet temperature: (900 8C-solid line, 950 8C-dashed line, 1000 8C-dash-dotted line) on rates of steam gasification (a), Boudouard (b), combustion (c), and WGS (d).

Figure 6 .

 6 Figure 6. Influence of O 2 concentration (0 vol %-solid line, 1.5 vol %-dashed line, 3 vol %-dash-dotted line) on char conversion (a), temperature profile (b), H 2 concentration (c), and CO concentration (d).

Figure 7 .

 7 Figure 7. Influence of O 2 concentration (0 vol %-solid line, 1.5 vol %-dashed line, 3 vol %-dash-dotted line) on rates of steam gasification (a), Boudouard (b), combustion (c), and WGS (d).

Figure 10b .

 10b Figure 10b. Thus, although CO production is favored by the increase in the initial CO 2 concentration from 5 to 10 vol %, a further increase in the latter from 10 to 20 vol % actually lowers the syngas quality, because the net increase of CO is insignificant (2 %) and H 2 production is further decreased by 27 %. These results are consistent with the experimental results of Butterman and Castaldi 6 who observed that introducing 0 to 5 vol % CO 2

Figure 8 .

 8 Figure 8. Influence of H 2 O concentration (5 vol %-solid line, 10 vol %-dashed line, 20 vol %-dash-dotted line) on char conversion (a), temperature profile (b), H 2 concentration (c), and CO concentration (d).

Figure 9 .

 9 Figure 9. Influence of H 2 O concentration (5 vol %-solid line, 10 vol %-dashed line, 20 vol %-dash-dotted line) on rates of steam gasification (a), Boudouard (b), combustion (c), and WGS (d).

Figure 10 .

 10 Figure 10. Influence of CO 2 concentration (5 vol %-solid line, 10 vol %-dashed line, 20 vol %-dash-dotted line) on char conversion (a), temperature profile (b), H 2 concentration (c), and CO concentration (d).

Figure 11 .

 11 Figure 11. Influence of CO 2 concentration (5 vol %-solid line, 10 vol %-dashed line, 20 vol %-dash-dotted line) on rates of steam gasification (a), Boudouard (b), combustion (c), and WGS (d).

Table 1 .

 1 Examples of Syngas End-Uses and Their Approximate H 2 /CO Ratio Requirements 6

	Syngas End-Use	H 2 /CO Ratio
	Solid oxide fuel cells (SOFC)	4.0-6.0
	Gas turbine combustion	2.5-4.0
	Fischer-Tropsch (diesel fuels)	1.5-3.0
	Fischer-Tropsch-Fe and Co-based	0.5-1.5
	catalyst process

Table 2 .

 2 Expression for Diffusion Coefficient of Gaseous Species as a Function of Temperature and Pressure (Chapaman-Enskog formula)[START_REF] Teixeira | Gaz eification de charbon de granules de bois: comportement thermochimique et m ecanique d'un lit fixe continu[END_REF] CO 2 , and O 2 ), as well as the char properties, porosity e p , particle thickness e p and reaction pre-exponential factor A i , by three functions that take the form[START_REF] Teixeira | Prediction of the gasification kinetics of a single wood char particle from a limited set of parameters[END_REF]

	Chemical Species j	Diffusion Coefficient D j;N2 (m 2 /s)
	N 2 CO 2 H 2 O CO H 2 O 2	1:39 Á 10 24 Á 1:013Á10 5 P 1:13 Á 10 24 Á 1:013Á10 5 P 2:11 Á 10 24 Á 1:013Á10 5 P 1:40 Á 10 24 Á 1:013Á10 5 P 5:11 Á 10 24 Á 1:013Á10 5 P 1:43 Á 10 24 Á 1:013Á10 5 P	À Á 3=2 Á T 1173 À Á 3=2 Á T 1173 À Á 3=2 Á T 1173 À Á 3=2 Á T 1173 À Á 3=2 Á T 1173 1173 Á T À Á 3=2

log 10 x i 5f T; P j ; e p ; e p ; A i À Á

  -21, and the coefficient values are listed in Table3.Combustion function log 10 x 1 5a1bT1cP O2 1de p 1ee p 1fA O2 1gT 2

	1hP O2	2 1ie p	2 1je p	2 1kTP O2 1lTe p 1mTA O2	(19)
	Boudouard function
	log 10 x 2 5a1bT1cP

CO2 1de p 1ee p 1fA CO2 1gT 2 1hP CO2 2 1iTe p 1jTe p 1kTA CO2 1le p e p 1me p A CO2 1ne p A CO2 (20) Steam gasification function log 10 x 3 5a1bT1cP H2O 1de p 1ee p 1fA H2O 1gT 2 1hP H2O 2 1iTe p 1jTe p 1kTA H2O 1lP H2O A H2O 1me p e p 1ne p A H2O 1oe p A H2O

Table 3 .

 3 Coefficient Values of the Functions for the Three Heterogeneous Reactions[START_REF] Teixeira | Prediction of the gasification kinetics of a single wood char particle from a limited set of parameters[END_REF]

	Reaction	Steam Gasification	Boudouard	Oxidation
	a b	22:38310 1 2 :92310 22	23:30310 1 4:18310 22	27:83 8:65310 23
	c	2.63	2.86	2:32310 1
	d	3 :72310 2	6:01310 2	22:89310 2
	e f g	24:33 3 :32310 26 21:02310 25	26:42 1:48310 28 21:45310 25	23:52 1:13310 210 23:12310 26
	h i j k	22:88 24:08310 21 3 :39310 23 22:59310 29	22:90 25:86310 21 4:86310 23 21:12310 211	26:59310 1 2:85310 4 3:78 24:95310 23
	l m n o	8 :38310 27 1 :19310 2 28:17310 25 1 :05310 26	1:34310 2 22:87310 27 4:04310 29	29:54310 22 28:36310 214

Table 4 .

 4 Expression for Production Term of Each Chemical Species

	Chemical Species j	Expression for Production Term

Table 5 .

 5 Expression of the Thermal Conductivity of Gaseous Species and Char as a Function of Temperature 10977ÁT 2 9.42549Á10 25 ÁT 2 1 8.05548Á10 28 ÁT 3 2 3.35367Á10 211 ÁT 4 1 5.15605Á10 215 ÁT 5 65-2500 37 O 2 21.7536 1 0.1224ÁT 2 1.322444Á10 24 ÁT 2 1 1.7804Á10 27 ÁT 3 21.200176Á10 210 ÁT 4 1 2.9817302Á10 214 ÁT 5 70-1500 37 CO 2 20.341914 1 0.0314ÁT 1 1.170458Á10 24 ÁT 2 2 1.281Á10 27 ÁT 3 1 5.7923Á10 211 ÁT 4 2 9.72044Á10 215 ÁT 5 190-2000 37 H 2 24.0803 1 0.9858ÁT 2 1.330466Á10 23 ÁT 2 1 1.1217Á10 26 ÁT 3 -3.25582Á10 210 ÁT 4

			Temperature
	Species j	Thermal Conductivity (mW/m/K)	Range (K)	Refs.
	N 2 CO H 2 O	20.3721 1 0.16-1500 80-2200 20.42832 1 0.09941ÁT -5.96573Á10 25 ÁT 2 1 3.81583Á10 28 ÁT 3 -1.43131Á10 211 ÁT 4 1 2.56748Á10 215 ÁT 5 148 1500	37 37 37
	C	95.8	307	38

Table 6 .

 6 Composition and Properties of Char from Maritime Pine Wood Chips and Pellets

		Van de Steene et al., 40	Teixeira et al., 24,33	Teixeira et al., 33
		Char Chips	Char Chips	Char Pellets
	Proximate analysis (wt % dry basis)
	Ash	1.4	1.7	1.4
	Volatile Matter	4.9	4.0	2.0
	Fixed carbon (by difference)	93.7	94.3	96.6
	Ultimate analysis (wt % dry basis)
	C	89.8	92.6	92.8
	H	2.2	1.0	1.3
	N	0.1	0.2	0.3
	O	6.1	3.8	3.1
	S LHV (MJ kg 21 dry basis) Char bed bulk density (g cm 23 ) Particle density (g cm 23 )	0.001 -a -a -a	<0.2 33.4 0.13 0.33	>0.2 33.1 0.37 0.66
	Particle porosity (-)	-a	0.74	0.51
	Particle average thickness (mm)	-a	5.2	4.2
	a

Table 7 .

 7 Operating Conditions at the Bed Inlet for Char Gasification Experiments

		Van de	Teixeira et al., 24,33
		Steene et al., 40	Char Chips
	Parameters	Char Chips	and Pellets
	Gas (%; mol/min)
	H 2 O	19; 2.35	28; 3.2
	CO 2	9; 1.12	8.2; 0.9
	O 2	2; 0.25	2.7; 0.3
	N 2	59; 7.31	61.1; 7.0
	CO	4; 0.5	0
	H 2	7; 0.87	0
	Solid char (g/min)	28	28
	Temperature (8C)	950	1028
	Total pressure (atm)	1.01	1
	Gas velocity (m/s)	0.68	0.7
	Bed height (cm)	65	65

Table 8 .

 8 Experimental and Predicted Results for Char Bed Outlet Conditions

		Van de
		Steene	Teixeira	Teixeira
		et al., 40 Char	et al., 24,33	et al., 33 Char
		Chips	Char Chips	Pellets
		Exp. Num. Exp. Num.	Exp.	Num.
	T (8C)	700	688	770	731	730	692
	H 2 O (vol %)	9.8	8.2	11.5	11.2	11.5	9.5
	CO 2 (vol %)	10.8	10.2	10.0	11.0	9.0	11.1
	O 2 (vol %)	0.6	0.0	0.0	0.0	0.0	0.0
	H 2 (vol %)	15.9	15.5	14.0	13.4	13.5	14.8
	CO (vol %)	9.8	12.4	11.2	10.5	11.5	11.4
	N 2 (vol %) Solid char (g/min)	52.9 3.1	53.7 3.6 $2.5 a 53.3	53.9 2.3 $2.8 b 54.6	53.1 2.9
	Char conversion,	89	71	95	87	95.0	99.0
	X
	a Calculated at a bed height of 50 cm.

b Calculated at a bed height of 15 cm.

Table 9 .

 9 Kinetic Parameters Used in the Model for the Water Gas Shift Reaction by Teixeira et al.[START_REF] Teixeira | Gasification of continuous wood char bed: modelling and experimental approach[END_REF]

		k WGS	K eq	WGS
	A WGS Ea WGS	2.78 m 3 mol 21 s 21 1510.70 J mol 21	2.65 3 10 22 32911 J mol 21

Table 10 .

 10 Operating Conditions for Sensitivity Analysis

		O 2	H 2 O	CO 2	N 2
	T (8C)	(vol %)	(vol %)	(vol %)	(vol %)

Table 11 .

 11 Summary of Simulation Results from Sensitivity Analysis

		Char
		conversion,	Temperature,	H 2	CO	H 2 /
		X (%)	T (8C)	(vol %)	(vol %)	CO
	T (900-1000 8C)	127%	114	117%	133%	20.1
	O 2 (0-3 vol %)	148%	186	120%	1133%	20.4
	H 2 O (5-20 vol %)	117%	26	1150%	233%	11.0
	CO 2 (5-10 vol %)	123%	21.7	212%	18%	20.1
	CO 2 (10-20 vol %)	141%	128	227%	12%	20.2

  2 / CO ratio. O 2 and H 2 O have the strongest influence on syngas CO and H 2 concentrations respectively. A small increase in O 2 concentration from 0 to 3 vol % increases char conversion and bed temperature to produce a syngas rich in CO (low H 2 /CO ratio) and thus suitable for FT synthesis. An increase in H 2 O concentration from 5 to 20 vol % results in a significant increase in H 2 (high H 2 /CO ratio) and thus favors the production of syngas for use in gas turbines and solid oxide fuel cells (SOFC).

[Color figure can be viewed in the online issue, which is available at wileyonlinelibrary.com.]

Acknowledgments

This work was funded by Suez Environnement and the French "Investments for the future" Program managed by the National Agency for Research under contract ANR-10-LABX-22-01. The authors gratefully acknowledge these supports.

Notation


Doan Pham
email: doan.phamminh@mines-albi.fr

Ange Nzihou

Patrick Sharrock

Carbonated hydroxyapatite starting from calcite and different orthophosphates under moderate hydrothermal conditions: Synthesis and surface reactivity in simulated body fluid

Keywords:

may come

Carbonated hydroxyapatite starting from calcite and different orthophosphates under moderate hydrothermal conditions: Synthesis and surface reactivity in simulated body fluid

Introduction

The inorganic mineral phase of human teeth and bone is mainly composed of calcium hydroxyapatite (Ca-HA, Ca 10 (PO 4 ) 6 (OH) 2 ), with different substitutions such as sodium, magnesium, chloride, fluoride, carbonate, etc. [1][2][START_REF] Bilezikian | Principles of Bone Biology[END_REF]. This explains the interest in the development of Ca-HA and substituted Ca-HA in the biomaterials field, due to their biocompatibility and bioactivity [START_REF] Thian | [END_REF]5]. Substituted Ca-HA is formed when calcium, phosphate, and/or hydroxyl groups of Ca-HA structure are replaced by other metals and functional groups. Thus, bone mineral has usually the molar ratio of calcium to phosphate varying in the range of 1.37-1.87 compared with that of 1.67 for stoichiometric Ca-HA [START_REF] Bilezikian | Principles of Bone Biology[END_REF]. In general, carbonate is found to be the most abundant substitution and its content in bone mineral varies in the range of 3-8 wt.% [6]. Carbonate groups can insert in the apatitic structure by replacing phosphate and/or hydroxyl groups to form carbonated hydroxyapatite (CAP). Conventionally, CAP is A-type CAP when carbonate groups replace phosphate groups; B-type CAP when carbonate groups replace hydroxyl groups; and A-B-type CAP when carbonate groups replace simultaneously phosphate and hydroxyl groups.

The synthesis of CAP calls usually for the classical wet and dry methods. In the wet process, soluble calcium and phosphates are precipitated in an aqueous solution in the presence of carbonate anions [7,8]. This process leads generally to the formation of B-type CAP [7] but A-B-type CAP may be also formed [9]. In the dry process, A-type CAP could be generated by heating pure Ca-HA under CO 2 flux at high temperature in order to replace hydroxyl groups by carbonate groups [10]. CAP could be also obtained by the common mechano-chemical process as demonstrated by Suchanek et al. [11], wherein an aqueous suspension of Ca(OH) 2 and CaCO 3 (or Na 2 CO 3 ) and (NH 4 ) 2 HPO 4 are treated in a laboratoryscale mill. But this last way requires severe reaction conditions.

We recently developed a new one-step synthesis process of Ca-HA from CaCO 3 and orthophosphoric acid (H 3 PO 4 ) as economical starting materials at moderate synthesis conditions [12]. The advantage of this process compared to classical Ca-HA synthesis pathways remains in the fact that pure Ca-HA can be obtained without any further purification step, thanks to the absence of all counter ions in the reaction medium [12]. A modification of this process was then developed for the one-step synthesis of CAP under moderate hydrothermal conditions [13]. CAP containing 4.2 wt.% of carbonate could be obtained from the reaction of CaCO 3 and H 3 PO 4 in a close batch reactor under CO 2 pressure. CaCO 3 was chosen because it is the source of both calcium cations and carbonate anions for the formation of CAP, and also because it has considerable industrial importance and abundant diversity as a biomineral [START_REF] Mozafari | Challenges of spontaneous formation of calcite during bioactivity assessment on bioactive materials interfaces[END_REF]. In this work, we firstly compared the reaction of calcite with different alkali orthophosphates and orthophosphoric acid in order to select the best orthophosphate. Then, we investigated the influence of different parameters including the content of water in the initial suspension of CaCO 3 , the reaction temperature and the reaction time, on the content of carbonate inserted in the final solid CAP. The in vitro biomineralization activity of the solid CAP product containing the highest carbonate content was also investigated before and after thermal treatment at different temperatures.

Materials and methods

Orthophosphoric acid (H 3 PO 4 , 85 wt.% in water) was purchased from Merck. Calcite powder (CaCO 3 , d 50 of 23.6 mm), and alkali orthophosphates (NaH 2 PO 4 , KH 2 PO 4 , NH 4 H 2 PO 4 ) of analytical purity from Fisher Scientific were used as received without further purification. The reaction of CaCO 3 with these orthophosphates was carried out in a 250 mL stainless steel close batch reactor (Top Industry, France) which was equipped with an electrical heating jacket and a magnetic stirrer. When orthophosphoric acid was used, 10 g of calcite powder and 45-135 g of water were fed into the reactor. The reactor was closed and then 6.92 g H 3 PO 4 (85 wt.%) was quickly injected into the reactor via a syringe. When alkali orthophosphates were used, 10 g of water was set into the reactor. This last one was closed and 135 g of an aqueous solution containing 60 mmol of an alkali orthophosphate was quickly injected into the reactor. In all case, the molar ratio of calcium to phosphorus in the initial reaction mixture was 1.67. The reaction mixture was kept at 60 or 80 C under stirring of 800 rpm for 48-168 h of reaction. The conditions for the synthesis of different CAP were summarized in Table 1. During the reaction, the pressure in the batch reactor was measured by a manometer. When the reaction reached the desired reaction time, the reactor was freely cooled down to room temperature. Solid products were separated from liquid phase by filtration using 0.45 mm filter paper. Then they were air-dried at 50 C for 48 h before further analyses and characterizations.

Elemental analysis was carried out by inductively coupled plasma atomic emission spectroscopy (ICP-AES, HORIBA Jobin Yvon Ultima 2 apparatus). Infrared spectroscopy (IR) was carried out using a Shimadzu FTIR 8400S spectrometer in the wavenumber range 4000-500 cm À1 . X-ray diffraction (XRD) measurements were carried out using a Phillips Panalytical X'pert Pro MPD diffractometer. Scanning electron microscopy (SEM) observation was performed on a Philips XL30 ESEM apparatus. Thermogravimetry analysis (TG) was carried out with a SDTQ600 analyzer (TA Instruments) under air flow (100 mL min À1 ) with the heating rate of 20 C min À1 . The amount of remaining CaCO 3 in the final solid was calculated corresponding to the observed weight loss of CO 2 from 610 to 720 C.

Results

Orthophosphate comparison

Fig. 1 presents XRD results of the solid products obtained from the reaction of calcite with four orthophosphates under similar reaction conditions. In all cases, Ca-HA appeared as the main crystalline calcium phosphate. On the other hand, the decomposition of calcite was different from each other orthophosphate. Under the reaction conditions used, orthophosphoric acid allowed decomposing completely the introduced calcite, because the signal of the main diffraction peak of calcite at 29.4 was negligible (CAP-H pattern). For three alkali orthophosphates used, diffraction peaks of remaining calcite were clearly observed, which indicates that the reaction of calcite and these orthophosphates was not completed. This result was confirmed by TG analysis (results not shown). The remaining calcite content present in the final solids was 4.4, 1.4, 0.5 and <0.06 wt.% for CAP-Na, CAP-K, CAP-N and CAP-H, respectively. Probably, orthophosphoric acid with higher acidity compared to other alkali orthophosphates was more favorable for the dissolution of calcite particles, which is crucial for the formation of calcium phosphate phases [START_REF] Pham Minh | [END_REF]. When calcite particles were covered by solid calcium phosphate layers, the contact between these covered calcite particles and orthophosphates sources must be prevented and so the reaction is stopped or slowed down [START_REF] Pham Minh | [END_REF]16].

In addition to this higher performance for the decomposition of calcite, no washing step was needed for the purification of the solid product starting from orthophosphoric acid thanks to the absence of counter-ions such as alkalis. So for the next parametric study, CAP synthesis was performed only with orthophosphoric acid.

CAP from calcite and orthophosphoric acid -XRD characterization

Fig. 2 presents XRD patterns of the solid products obtained at different synthesis conditions. Both products obtained at 60 C (T60/t48/D4.5 and T60/t48/D9) contained octocalcium phosphate (OCP, Ca 8 (HPO 4 ) 2 (PO 4 ) 4 Á5H 2 O) as the main crystalline calcium phosphate, and high contents of remaining calcite. So the temperature of 60 C was not enough for the complete dissolution of calcite particles. Probably this temperature led to the formation of core-shell particles in which calcite core was covered by solid calcium phosphate layers (shell) [17]. This structure prevented the contact of remaining calcite core with phosphate species and slowed down or stopped the reaction at this temperature. The increase of the dilution of calcite suspension did not improve the dissolution of calcite at 60 C.

On the other hand, when the reaction temperature rose to 80 C, calcite could be completely dissolved as indicated its very weak signal of the principal XRD peak at 29.4 . At this temperature and for the same reaction time of 48 h, the dilution of CaCO 3 suspension had no remarkable influence on the formed product. The main crystalline calcium phosphate phase identified for the products synthesized at 80 C was Ca-HA. However, CAP (JCPDS standard No. 96-900-3550, 96-900-3551, 96-900-3552) might be also present in these products but their XRD diffractions are usually superposed on those of Ca-HA. So the presence of CAP needs to be confirmed by other analyses.

CAP from calcite and orthophosphoric acid -IR characterization

IR spectra of all the solid products are presented in Fig. 3. In all cases, characteristic peaks of phosphate groups could be found in the wavelength ranges of 650-550 cm À1 and 1300-910 cm À1 [13]. For two solids obtained at 60 C, no signal of carbonate groups was observed, despite remarkable contents of remaining calcite as shown by XRD analysis (Fig. 2). As mentioned above, calcite remnants must be covered by calcium phosphate layers and therefore could not be detected by diffuse reflectance IR analysis, which records only surface information for solid samples.

For the solids synthesized at 80 C, absorption bands of carbonate groups could be clearly observed at 1545, 1450, 1415, 880, and 870 cm À1 . All these absorption bands could be attributed to carbonate groups inserted in the apatite structure of Ca-HA [ [START_REF] Elliott | Studies in Inorganic Chemistry 18: Structure and Chemistry of the Apatites and Other Calcium Orthophosphates[END_REF][START_REF] Fleet | [END_REF][20]. Carbonate groups replaced both phosphate and hydroxyl groups to form A-B-type CAP, as illustrated by their characteristic absorption bands at 1545 and 880 cm À1 (A-type) and 1450, 1415 and 870 cm À1 (B-type) [START_REF] Elliott | Studies in Inorganic Chemistry 18: Structure and Chemistry of the Apatites and Other Calcium Orthophosphates[END_REF][START_REF] Fleet | [END_REF]. The elemental analysis of three solid products obtained at 80 C showed that the molar ratio of calcium to phosphorus varied in 1.83-1.90 range, which was higher than the value of the stoichiometric Ca-HA. This can be explained by the insertion of carbonate groups at the position of phosphate groups.

CAP from calcite and orthophosphoric acid -TG characterization

For the use as biomaterials, calcium phosphate based products are usually treated by a sintering step [21]. Thermal behavior of the prepared solids was investigated in large range of temperature as shown in Figs. 4 and5 . For the solid prepared at 60 C (Fig. 4), after the first weight loss about 100 C due to the removal of surface humidity, different weight losses in the temperature range of 100-600 C were observed for the dehydration of OCP [START_REF] Elliott | Studies in Inorganic Chemistry 18: Structure and Chemistry of the Apatites and Other Calcium Orthophosphates[END_REF][22][23][24][25], which was the main calcium phosphate present in these solids. Remaining calcite showed an intense weight loss at about 610 C, followed by the decarbonation of CAP at 750 C. No further weight loss was observed up to 1200 C for these two solids.

For the solids synthesized at 80 C, after the dehydration of surface moisture around 100 C, the dehydration of hydrogenophosphate groups took place in the range of 100-600 C. These weight losses were much lower compared to the solids synthesized at 60 C (Fig. 4). As expected, no notable signal of calcite decomposition was detected around 610 C. On the other hand, the decarbonation of CAP took place with important weight losses. This decarbonation started at about 750 C and finished at about 1200 C. The last weight loss around 1250 C was due to the dehydration of hydroxyl group of apatitic structure to form oxyhydroxyapatite [5]. So, when the synthesis temperature was set at 80 C, the reaction time and the dilution of the initial calcite suspension had no notable influence on the final CAP. Three solid products synthesized at 80 C seemed also to have similar thermal behavior. From TG analysis, the content of carbonate inserted in the apatitic structure could be calculated, supposing that weight loss in the temperature range of 750-1250 was due to the thermal decarbonation. Thus, three solid products synthesized at 80 C from calcite and orthophosphoric acid had the carbonate content in the range of 4.2-4.6 wt.%.

CAP from calcite and orthophosphoric acidthermal treatment and biomineralization activity in SBF solution

use as biomaterials, hydroxyapatite based products, including CAP, may need stabilization step by thermal treatment. In the case of CAP, the thermal treatment should not lead to deep decarbonation, in order to maintain apatitic carbonate content similar to that of bone mineral [6]. In stability of CAP synthesized 80 C (T80/t48/D4.5) is discussed. This T80/t48/D4.5 was chosen because of its lowest content of remaining CaCO 3 . The reactivity of solid products in SBF test, as the biomineralization activity, is also reported. This test was carried out with T80/t48/D4.5 thermally treated at the temperature range of 105-1400 C under air atmosphere.

Thermal treatment of T80/t48/D4.5 powder was carried out in the TG-DSC analyzer under air atmosphere. The powder was heated at 10 C min À1 up to desired final temperature and was kept at this temperature for 5 h. The final temperature varied in the range 600-1400 C. After, the oven was freely cooled down to room temperature. IR spectra were rapidly recorded on sintered products, and are presented in Fig. 6. Although the decarbonation of CAP started around 750 C (TG analysis, Fig. 5), the thermal treatment in the temperature range of 600-800 C seemed to have no impact on the carbonate content of T80/t48/D4.5 solid. This may be explained by a low kinetic decarbonation of T80/t48/D4.5 at 600-800 C. When the sintering temperature rose to 1000 C, a partial decarbonation could be observed, leading to reduction of apatitic carbonate groups. At 1200-1400 C, absorption bands of carbonate groups disappeared, indicating that the decarbonation was complete. The calcination at 1200 C led to the characteristic IR spectra of Ca-HA, while the calcination at 1400 C decomposed Ca-HA into tetracalcium phosphate (TTCP) and tricalcium phosphate (TCP) [START_REF] Elliott | Studies in Inorganic Chemistry 18: Structure and Chemistry of the Apatites and Other Calcium Orthophosphates[END_REF]26]. According to previous reports [START_REF] Elliott | Studies in Inorganic Chemistry 18: Structure and Chemistry of the Apatites and Other Calcium Orthophosphates[END_REF][START_REF] Fleet | [END_REF]26], the thermal treatment under air atmosphere is expected to increase the crystallinities of Ca-HA and CAP, and decrease the apatitic carbonate content of CAP, as confirmed by IR results (Fig. 6). But at the very high temperature of 1200-1400 C, Ca-HA and CAP decomposed into TCP and TTCP.

Biomineralization activity of sintered solids was investigated using simulated body fluid (SBF) solution. Synthetic SBF solution was prepared according to previous work [27]. For this test, a solid sintered at a specific temperature was submerged into SBF solution and the obtained suspension was thermostated at 37 C for 7 days. SEM study was then performed on the powders recovered after rising with water and drying at room temperature.

On the left-hand-side of Fig. 7, the sintering at 600-1000 C had no significant influence on the apparent morphology of CAP particles. Particles of flat-needle-like morphology dominated CAP powders, dried at 105 C or sintered at 600-1000 C. When the temperature rose to 1200-1400 C (left-hand-side, Fig. 8), sintering took place with the increase of particle size, true density and the decrease of porosity and specific surface area [28,29].

For CAP powders treated at 105-1000 C, despite the decrease of carbonate content when the calcination temperature increased as showed by TG analysis, CAP particles were covered by new calcium phosphate layers after contact with SBF solution (Fig. 7, right-hand-side). Previous reports showed that Ca-HA based materials had appropriate biomineralization activity for the growth of calcium phosphate layers in SBF solution [27,30,31]. SBF solution was also found as favorable medium for the precipitation of calcium and phosphate to form Ca-HA [32,33]. For our solid synthesized at 80 C, the presence of trace amount of remained calcite must not influence the growth of Ca-HA because calcite has been shown as biomineral and its growth took place in SBF solution [START_REF] Mozafari | Challenges of spontaneous formation of calcite during bioactivity assessment on bioactive materials interfaces[END_REF]33]. For the solids calcined at 1200-1400 C, although the thermal decarbonation was complete, the growth of new calcium phosphate layers also took place (Fig. 8, right-hand-side). In fact, Ca-HA, and non-apatitic calcium phosphates including TTCP and TCP, which resulted from thermal treatment of CAP at 1200-1400 C, are also known as active biomineralization materials in SBF solution [26,31,34,35].

Conclusions

The reaction of calcite and different orthophosphates was investigated. Alkali phosphates did not allow a complete decomposition of calcite at 80 C in a close batch reactor.

On the other hand, orthophosphoric acid led to the complete decomposition of calcite powder, and carbonated hydroxyapatite (CAP) was formed. The reaction temperature of 80 C was found to be crucial for the total decomposition of calcite and thus the formation of apatitic phase. The reaction time and the dilution of calcite suspension had no significant influence on the reaction.

CAP powder starting from orthophosphoric acid and calcite was thermally stable at 600 and 800 C under air atmosphere. At 1000 C for 5 h, partial decarbonation was observed. The decarbonation was total at 1200 or above. All CAP and its thermal decomposition products were found to be reactive in SBF solution, promoting crystallization of biomimetic apatites.

Fig. 2 .Fig. 3 .

 23 Fig. 2. XRD patterns of the solid products obtained at different synthesis conditions; (&) Ca-HA (JCPDS standard No. 01-072-1243); (~) OCP (JCPDS standard No. 00-026-1056); (*) calcite (JCPDS standard No. 00-047-1743).

Fig. 4 .Fig. 5 .

 45 Fig. 4. TG analysis of the solid products synthesized at 60 C.

Fig. 6 .

 6 Fig.6. IR spectra of T80/t48/D4.5 after thermal treatment at different temperatures.

Fig. 7 .

 7 Fig. 7. SEM images of T80/t48/D4.5 treated at 105-1000 C before (left-hand-side) and after (right-hand-side) contact with SBF solution.

Fig. 8 .

 8 Fig. 8. SEM images of T80/t48/D4.5 treated at 1200-1400 C before (left-hand-side) and after (right-hand-side) contact with SBF solution.

Table 1

 1 Synthesis conditions of CAP from CaCO 3 (100 mmol or 10 g), alkali orthophosphates, and H 3 PO 4 (66.7 mmol); D is the mass ratio of water to calcite in the initial suspension. The final pressure in the batch reactor varied in the range of 9-16 bar.

	Orthophosphate	T	t	D	Water	Solid product	Note
		( C)	(h)		(g)	designation
	NaH 2 PO 4	80	48 13.5 135	CAP-Na	Comparison of
	KH 2 PO 4	80	48 13.5 135	CAP-K	orthophosphate
	NH 4 H 2 PO 4	80	48 13.5 135	CAP-N
	H 3 PO 4	80	48 13.5 135	CAP-H
	H 3 PO 4	60	48 4.5 45	T60/t48/D4.5 Parametric
	H 3 PO 4	60	48 9	90	T60/t48/D9	study
	H 3 PO 4	80	48 4.5 45	T80/t48/D4.5
	H 3 PO 4	80	48 9	90	T80/t48/D9
	H 3 PO 4	80	168 9	90	T80/t168/D9

Fig. 1. XRD patterns of the solid products obtained from the reaction of calcite with different orthophosphates at 80 C in close batch reactor.

Acknowledgment

The authors gratefully acknowledge colleagues at RAPSODEE Center for technical help.