Keywords: VOC, Recuperative incinerators, Combustion, Turbulence, Emissions

can be drastically reduced. In the best case and in spite of the fact that the turbulator geometry is not yet optimised, CO emissions are reduced from 1990 ppmv down to 600 ppmv. This is equivalent to a possible decrease in the exit temperature of 30 °C keeping CO emissions at the same level in the case where the turbulator is used. This is also equivalent to a reduction of 5.4% in the NG consumption of the incinerator.

Improving VOC recuperative incinerators performances by increasing turbulence levels inside the combustion chamber-experimental results S. Salvador

Abstract

Recuperative incinerators are used to eliminate volatile organic compounds (VOC) present in air fluxes prior to their release into the atmosphere. In this process, the air flow is heated to approximately 700-800 °C and maintained at this temperature for roughly 0.5-1 s. VOCs are cracked into light hydrocarbon species, and oxidised with the oxygen of the air flow into CO 2 and H 2 O. To reach this temperature, it is necessary to feed the incinerator with external Natural Gas (NG). With the goal of minimising the NG consumption of the process, the reaction temperature must be maintained at a value as low as possible. Under these conditions, the factor that limits the incinerator performance is CO emissions, that can rapidly exceed several tens of thousands of ppmv if the temperature is decreased; the upper limit regarding regulation is 100 mg m À3 at STP, or 80 ppmv. Turbulators with various geometry can be introduced inside the combustion chamber (CC) to enhance the CO oxidation to CO 2 reaction, through an increase in the turbulence level of the flow. Nevertheless, no quantitative data about the impact of the technique on the combustion can be found in the literature.

In the present paper, experimental results are reported on the effect of inserting turbulators of different configuration inside the combustion chamber. Depending on the position of the turbulator, CO emissions

Introduction

Very significant air flows containing volatile organic compounds (VOC) are generated in numerous industrial sectors across the world. On the planetary scale, natural sources of VOC and in particular vegetation and forest fires prevail: they account for 90% of the emissions. In industrialised areas, these emission sources become minor [START_REF] De | Elimination de compose ´s organiques volatils (COV)-Oxydation catalytique de lÕO-Xyle `ne sur catalyseurs palladium et platine/Ze ´olithes[END_REF]. In France, 16% of the VOC emissions are of natural origin while 84% are related to human activities [START_REF] Dueso | La re ´duction des e ´missions de compose ´s organiques volatils dans lÕindustrie[END_REF]. With respectively 46% and 31% of the anthropic emissions, the transport and the use of organic solvents are the principal sources of VOC [START_REF] Dueso | La re ´duction des e ´missions de compose ´s organiques volatils dans lÕindustrie[END_REF]. It is within the framework of the VOC elimination resulting from the industrial production that our work is aimed. Since VOC are harmful to animals and plants, and additionally contribute to the green house effect, it is necessary to treat these air streams in order to remove VOCs prior to their release into the atmosphere. One of the most commonly used techniques today is to oxidise them into CO 2 and H 2 O using a recuperative incinerator [START_REF] Ruddy | Select the best VOC control strategy[END_REF][START_REF] Van Der | Bagely-Pride, Thermal and catalytic incinerators for the controls of volatile organic compounds[END_REF]. The general principle of the process is schematised in Fig. 1. The air flux containing VOC is first preheated by crossing a gas-gas heat exchanger. It is then fed to a combustion chamber (CC). As shown in Fig. 2, this chamber is fitted at the inlet with an ''in vein'' burner supplied with Natural Gas (NG). Part of the air flow passes through perforated planes called ''wings'' and feeds the flame which is stabilised downstream of the burner. Two diaphragms are placed symmetrically above and below the burner wings, and leave two free sections (air veins), through which the main part of the air flux is blown. The diaphragm is employed [START_REF] Kara | Oxydation thermique des Compose ´s Organiques Volatils (COV) en e ´purateur re ´cupe ´ratif-Optimisation des outils de dimensionnement[END_REF]:

• to create air recirculation downstream, in order to obtain good mixing between the hot gases from the flame and the ''cold'' air flow that crosses the air veins, • to increase the turbulence levels inside the combustion chamber, in order to enhance the oxidation of the VOCs.

The aperture of the diaphragms can be adjusted. This will also affect the repartition of the air flow between the burner and the air veins, and consequently control the richness of the NG flame.

The temperature level inside such a reactor can be characterised by the output temperature of the CC, where the temperature field is quite uniform on a cross section. This level is typically in the range of 700-800 °C. This hot air stream crosses the heat exchanger, and is cooled down to typically 350-400 °C before being released to the atmosphere.

One of the advantages of this technology is that, by using large width burners and combining them into H or I shaped ramps, it is possible to build incinerators with very high treatment capacities [START_REF] Kara | Oxydation thermique des Compose ´s Organiques Volatils (COV) en e ´purateur re ´cupe ´ratif-Optimisation des outils de dimensionnement[END_REF].

The environmental performance of a VOC incinerator is evaluated through the following criteria:

• the destruction efficiency of VOCs. Depending on the VOC toxicity, efficiencies ranging between 95% and 99.999% can be aimed at [START_REF] Bussac | Veille technologique sur le traitement des COV[END_REF][START_REF] Mortgat | Traitement des COV-Les proce ´de ´s dÕincine ´ration pour anticiper sur le futur bilan solvant[END_REF]. • NO x emissions. The flame generated by the burner also has the role to stabilise the combustion inside the CC. As a consequence, this flame has to be ''hard'', and its temperature is typically between 1950 °C and 2150 °C: thermal NO x are formed by the flame [START_REF] Kara | Oxydation thermique des Compose ´s Organiques Volatils (COV) en e ´purateur re ´cupe ´ratif-Optimisation des outils de dimensionnement[END_REF][START_REF] Stansel | CO and NO x emissions from controlled-air burner: Experimental measurements and exhaust correlations[END_REF]. Todays regulation sets a limit at 100 mg m À3 at STP of equivalent NO 2 , or 50 ppmv NO x (NO + NO 2 ) at the exit of the incinerator [START_REF]Arre ˆte ´relatif aux pre ´le `vements et a `la consommation dÕeau ainsi quÕaux e ´missions de toute nature des installations classe ´es pour la protection de lÕenvironnement soumises a `autorisation, Ministe `re de lÕame ´nagement du territoire et de lÕenvironnement[END_REF]. • CO emissions. The species CO is an intermediate species in the combustion of VOCs. The reaction mechanism to oxidise it into CO 2 is quite slow. For low temperature operation, the concentration of CO at the exit of the incinerator can increase to more than several thousands ppmv [START_REF] Kara | Oxydation thermique des Compose ´s Organiques Volatils (COV) en e ´purateur re ´cupe ´ratif-Optimisation des outils de dimensionnement[END_REF].

In a previous work [START_REF] Bussac | Veille technologique sur le traitement des COV[END_REF], we have shown that the last criterion, e.g. CO emissions, is the principal factor that does not allow one to reduce the temperature inside the CC, and consequently does not allow one to reduce the NG consumption. In this paper, we show that increasing the turbulence level inside the CC accelerates the oxidation of CO into CO 2 . We first describe the incinerator used for the experiments, as well as the two turbulators that were developed and tested. Then we examine the impact of these turbulators on the combustion process, in terms of their position along the CC.

Experimental setup

A prototype of an industrial recuperative incinerator at half scale was designed and built at Ecole des Mines dÕAlbi-Carmaux. A detailed description of the apparatus may be found in [START_REF] Kara | Oxydation thermique des Compose ´s Organiques Volatils (COV) en e ´purateur re ´cupe ´ratif-Optimisation des outils de dimensionnement[END_REF]. Experimental work on NO emissions reduction conducted on this apparatus has been described elsewhere [START_REF] Salvador | Reduction of NO emissions from a VOC recuperative incinerator by dilution of the fuel supply[END_REF]. A general view of the incinerator and surroundings is provided in Fig. 3. The cross section of the CC is 320 mm wide and 360 mm high. The 75 kW burner consists of a 300 mm large and 80 mm high ramp. Tests were carried out with a diaphragm height of 109 mm, resulting in air veins height of 31 mm. It was estimated, from numerical modelling [START_REF] Kara | Oxydation thermique des Compose ´s Organiques Volatils (COV) en e ´purateur re ´cupe ´ratif-Optimisation des outils de dimensionnement[END_REF], that approximately 18% of the air flow crosses the wings of the burner. The air flux velocity upstream of the burner was homogenised on a cross section using three perforated plates (Fig. 2, flow straightener). The length of the CC downstream of the burner is 2700 mm. A refractory steel tube heat exchanger is used to recover part of the energy of the flue gas. A 25 kW electric heater is placed downstream to accurately set the air preheating temperature before entering the CC. Liquid VOC, ethyl acetate in this case, is injected by spraying in compressed air downstream of the electric air heater. The air flow, the liquid VOC flow rate and the air preheating temperature were measured and controlled. Experiments were conducted several times so to ensure consistent results. For instance NO x mea-Fig. 3. General view of the recuperative incinerator and of its surroundings.

surements or CO measurements, if repeated on different days, would lead to values of 20 ± 2 ppmv and 1990 ± 50 ppmv respectively.

The conditions for all the tests reported here are the following. The air flow rate was set to 468 m 3 s À1 (at STP), and the VOC concentration was maintained at 9.1 g m À3 (at STP). The air preheating temperature was set to 376 °C and the exit temperature was fixed at 715 °C.

The first turbulator is designed according to the following characteristics; its geometrical arrangement is illustrated in Fig. 4. The diameter of the tubes must be large enough so that the tubes do not bend under their weight; 17.1 mm o.d. refractory steel (304 L) tubes are used. Since the average velocity for the hot air is approximately 4.5 m.s À1 , the Reynolds number for each tube is 3100. For this flow regime, i.e. for Reynolds numbers ranging from 60 to 5000, vortices are formed downstream of each tube and are blown downstream. It is expected that this regime will promote turbulence in an important volume downstream of each tube.

We have adopted here, as a single assembly module of tubes, 2 vertical staggered rows of 6 tubes each. The staggered assembly is known as a classical way to maximise the exchange between a fluid and the external surface of a ''bank'' of tubes, while minimising the pressure drop in the fluid phase. This is why this configuration was adopted here. The 12 tubes have a projected surface of 65 664 mm 2 , which represents an obstructed surface of 57% of the section of the CC.

From the results of a numerical model developed in a previous work [START_REF] Kara | Oxydation thermique des Compose ´s Organiques Volatils (COV) en e ´purateur re ´cupe ´ratif-Optimisation des outils de dimensionnement[END_REF], it is thought that the region where the oxidation of CO is limited by micro-mixing efficiency is located around an horizontal plane at half the height of the CC. This is why no tubes were placed close to the upper and bottom walls of the CC.

In an industrial application, a number of these assembly can be placed inside the CC. In this paper, the case of one assembly and the case of two assemblies have been examined.

Results

Flue gases at the exit of the CC are sampled using a water cooled probe. They are analysed in real time for O 2 using a paramagnetic analyser (SERVOMEX), for CO and NO x using non-dispersive Infra Red analysers (SERVOMEX), and for total hydrocarbons (unburned hydrocarbons) using a Flame Ionisation Detector (COSMA S.a. Environnement). The oxygen content of the flue gases was found to be 17.4%.

The turbulator was placed closed to the burner, at 500 mm, and fitted with a steel wire that crosses the end wall of the CC. The incinerator was started and the steady state was reached after 5 h. The turbulator was then moved step by step by pulling on the wire from the outside and measurements were taken once the steady state regime was re-established.

We report in Fig. 5 the evolution of CO concentration at the exit of the CC versus the position of the turbulator. The CO emissions without the turbulator were recorded at 1990 ppm. This value is recovered when the turbulator is placed close to the CC exit, where it has no more impact. It is clear from this figure that the turbulator can enhance the oxidation of CO into CO 2 . The impact is maximum when the turbulator is placed 500 mm downstream the burner. The effect decreases when the turbulator is moved further from the burner, as indicated by the increase in CO emissions. Positions closer to the burner were not tested because we have estimated that, regarding the high temperatures reached inside the flame zone, the life time of the turbulator, both during our tests and in future industrial applications, would be severely shortened. These results indicate that there is a large zone of the CC where the reaction rate for the oxidation of CO into CO 2 depends on the turbulence intensity.

It is clear today that a gas phase reaction such as the ones involved in this combustion process can be controlled by two types of mechanism: (i) In the case where reactants and products are perfectly mixed and where the temperature is homogeneous, the reaction rate is determined by the kinetics of the chemical reaction. TodayÕs numerical models describe this through the calculation of an Arrhenius reaction rate. The main factors affecting the reaction rate are the local temperature and the species concentrations. (ii) In the case where the reaction kinetics is rapid, the reaction rate is controlled by the micromixing of the reactant and of the product species. Magnussen et al. [START_REF] Magnussen | On mathematical models of turbulent combustion with special emphasis on soot formation and combustion[END_REF] proposed a simple theory to calculate this reaction rate. They show that this mechanism can control the combustion in numerous applications, including the case of premixed combustion. In this situation, the main factors affecting the reaction rate are the kinetic energy k and its dissipation rate e. Following the concept of Spalding et al. [START_REF] Spalding | Mixing chemical reaction in steady confined turbulent flames[END_REF], the term k/e represents a characteristic time for the turbulent structures; the reverse of this term is the local micro-mixing efficiency.

The dependence of the reaction rate of the CO oxidation reaction observed here indicates that in almost all the CC the reaction is controlled by micro-mixing; indeed increasing the turbulence level accelerates the reaction. The micro-mixing considered here should not be confused with mixing at the scale of the reactor. The tubes here do not homogenise the temperature and species concentrations at the scale of the CC, but introduce turbulence that will enhance mixing of fresh gases and burned gases at a scale closed to molecular scale. In such a situation of micro-mixing controlled reaction, an important increase in the temperature is required for a small increase in the reaction rate.

It is difficult at this stage to evaluate the volume of the reactor downstream of the turbulators that is affected by them. One could imagine from the previous results that the turbulator, placed 500 mm downstream of the burner, generates turbulence along the length of the CC and enhances CO oxidation, and that moving it downstream reduces the volume where turbulence levels are increased. This would explain the decrease in CO concentration. Moreover, it is not possible to establish from these results if another increase in the turbulence level could further increase the CO oxidation efficiency. In order to bring some understanding, tests were conducted with 2 modules such as the one used before, placed side by side. The results obtained in terms of CO emissions as a function of the position of the turbulator along the CC are given in Fig. 6.

Overall, it is clear that doubling the number of tubes increases the CO oxidation efficiency. Therefore the reaction is still controlled by micro-mixing in the case of the single module turbulator.

Let us now consider the evolution of CO emissions as a function of the position of the turbulator. When the turbulator is placed closed to the burner, it has an important impact. The CO emissions remain almost constant as the turbulator is moved up to 1 m. Beyond 1 m, CO emissions increase; naturally, at the end of the CC, the turbulator becomes useless, and the 1990 ppmv CO concentration is recovered. The improvement in CO oxidation by the one module turbulator and by the two modules turbulator appear not to remain proportional as the turbulators are moved along the chamber. It is not possible to give explanations in terms of physical phenomena, regarding the non linearity of the thermal, fluid mechanics and chemical phenomena involved here. It is still difficult from these results to determine the volume of the CC downstream of the turbulator in which turbulence is significantly intensified. We can only say from these results that it does not cover the whole length of the CC, otherwise the efficiency would always decrease when the turbulator is moved from the burner to the exit of the CC.

As far as the pressure drop introduced by the turbulators is concerned, it was measured with two static pressure tappings placed at the walls upstream of the burner and at the exit of the CC. The pressure drop was about 20 Pa for the two modules turbulator. This can be compared, for reference, with the total pressure drop of the incinerator that is circa 640 Pa.

If one focuses on the VOCs destruction efficiency, we can refer to the Total Hydrocarbons emissions. For all the experiments performed in this work, their concentration remained smaller than 10 ppm, whether the turbulators were used or not. It can be thought that turbulators have increased the kinetics of the chemical reactions that crack the VOCs into lighter hydrocarbons and the reaction to oxidise the HC into CO, CO 2 and H 2 O.

As far as NO emissions are concerned, one debate whether if the turbulators enhance their formation or not. The emissions were recorded during all experiments. As indicated by the plots in Fig. 7, the NO x emissions were similar for all the experiments. This is probably due to the fact that the turbulators were always placed downstream of the flame. NO x are known to form inside the flame [START_REF] Kara | Oxydation thermique des Compose ´s Organiques Volatils (COV) en e ´purateur re ´cupe ´ratif-Optimisation des outils de dimensionnement[END_REF]: the turbulators did not have any impact on their formation.

The question of designing the best turbulator for an industrial incinerator is beyond the scope of this paper. It can nevertheless be discussed here. In this context, the criteria to evaluate the global improvement of the incinerator through the addition of a turbulator are:

• the reduction in CO emissions, resulting in a possible decrease in the operating temperature and consequently in a decrease in NG consumption, at constant CO emissions. • the cost of using a turbulator, that stems from:

-buying the refractory materials; -installing the turbulator; -eventual maintenance; -additional cost for air blowing due to the increase of the incinerator pressure drop. Fig. 6. Evolution of the CO volume fraction at the exit of the combustion chamber versus the position of (}) the one module turbulator and of (X) the two modules turbulator.

As far as the last point is concerned, the present results indicate that the pressure drop should not be the limiting criterion for turbulators amounting to several modules of 12 tubes each.

The present results also indicate that, in order to limit the investment cost for the refractory materials together with the maintenance cost, the turbulator should be placed far from the burner, but not further than 1 m. Should one consider the efficiency of more than 2 turbulator modules, additional tests are required. More generally, the tubes diameter, their spacing and the distance between each row should also be varied and optimised. Nevertheless, the large number of parameters entering in the design of the turbulator lead to a very large number of possible configurations, that could not be tested in the frame of this work.

Should numerical modelling be envisioned as a way to optimise the geometry of a turbulator, the difficulty will be to describe accurately the impact on the turbulence levels of the tubes. In the regime where vortices are periodically detached from each tube, and in the case of a complex geometry such as that of the incinerator, the CPU time cost will be very high.

Conclusions

When VOC recuperative incinerators are operated at moderate temperatures for reducing their natural gas consumption, the limitation encountered is not unburned VOC emissions but CO emissions.

It was shown that the introduction of a two row 12 tubes 17.1 mm turbulator inside the combustion chamber decreases the CO emissions from 1990 ppmv down to 780 ppm. The highest efficiency was obtained when the turbulator is placed 500 mm downstream to the burner.

Using two modules further decreases CO emissions down to 600 ppmv approximately. This variation in CO emissions is equivalent to a change in the CC temperature of 30 °C. This is equivalent to saving 5.4% of the NG supply. The ratio of the CO reduction efficiency of the two modules turbulator to that of the single module turbulator does not remain constant when the turbulator is moved along the CC. It was not possible to explain this at this stage. For a longer lifetime, a 24 tube turbulator should be placed far from the burner, but not further than 1 m from the burner for a good efficiency.

It is thought that the CO oxidation efficiency enhancement by turbulence level increase is due to an increase in the micro-mixing efficiency, that enables the chemical reaction acceleration through a better mixing, at a very small scale, of the burned gases with the fresh gases. The results presented here reveal that the incinerator performance is limited by micro-mixing if a turbulator is not used.

For all the configurations examined here, whether a turbulator is used or not, the emissions of hydrocarbons remained lower than 10 ppmv; it is likely that the use of a turbulator increases the VOC destruction efficiency.

NO x emissions are not affected by the presence of a turbulator. They remain constant approximately at 20 ppmv no matter what the design and location of the turbulator.

It is beyond the scope of this paper to determine an optimal geometry for the turbulator, regarding the large number of parameters entering in the design: diameter of the tubes, distance between two tubes in a row, distance between the rows, number of rows to take only the most common parameters. These results bring us to think that additional rows could still increase the CO oxidation efficiency. The pressure drop introduced by two modules is small: 20 Pa. Pressure drops should not be a limit in considering the addition of more than two tube rows.

Since the design of the turbulator proposed here is very simple, and the quantities of refractory materials required to build them are very small, this technique shows promise for industrial applications. It reduces the cost of VOC incineration through the decrease in NG consumption, and consequently limits the environmental impact of the process through CO 2 emissions reduction. NO x emissions are not affected, and VOC destruction efficiency is certainly enhanced.

Fig. 1 .

 1 Fig. 1. Principle of a VOC recuperative incinerator.

Fig. 2 .

 2 Fig. 2. General assembly of a VOC recuperative incinerator. (1) Burner; (2) diaphragms; (3) combustion chamber; (4) thermal insulation; (5) flow straightener; (6) air veins; (7) burner wings; (8) flame.

Fig. 4 .

 4 Fig. 4. Arrangement of the tubes in a single turbulator module.

Fig. 5 .

 5 Fig. 5. Evolution of the CO volume fraction at the exit of the combustion chamber versus the position of the one module turbulator.

Fig. 7 .

 7 Fig.7. Evolution of the NO x volume fraction at the exit of the combustion chamber versus the position of (}) the one module turbulator and of (X) the two modules turbulator.

  Ecole des Mines d'Albi-Carmaux, Centre Energe ´tique-Environnement, Laboratoire de Ge ´nie des Proce ´de ´s des Solides Divise ´s, UMR CNRS 2392, Campus Jarlard, Route de Teillet, 81013 Albi CT Cedex 09, France b Gaz de France Research and Development Division, Industrial Utilisation Section, 361 Avenue du Pre ´sident Wilson 93211 Saint-Denis La Plaine Cedex, France

a, * , Y. Kara b , J-M. Crussol c a c Sargi, 10, Cite´ d'Angouleˆme, 75011 Paris, France


J.-M Commandre
email: commandr@enstimac.fr

S Salvador

Lack of correlation between the properties of a petroleum coke and its behaviour during combustion

Keywords: Petcoke, NO emissions, SO 2 emissions

This work is an attempt to establish links between the properties of the petcoke that can be measured through standard analysis in the laboratory and the behaviour of the petcoke both in terms of its combustion velocity and the emission of gas pollutants.

A large number of petcokes, 22, taken from different sources, were burnt in an Entrained Flow Reactor (EFR) in conditions close to those of a fluidised bed of a cement plant precalciner. The burnout for the 22 petcokes ranged from 68% to 81%. The gas emissions resulting from their combustion were analysed. Prior to this, each type of petcoke was characterised in terms of its chemical composition, volatile matter and ash content, structural properties and LCV. The correlation between all these characteristics and the combustion velocity, SO 2 emissions and NO emissions in the EFR were systematically investigated.

The combustion velocity does not appear to be correlated to any of the properties. The emissions of SO 2 can be accurately predicted from the content in S of the petcoke. The emissions in NO are not directly linked to the N content of the petcoke. No simple correlation could be established to predict NO emissions.

Introduction

The refining of petroleum, during which the lighter fractions are extracted in the form of liquid or gaseous fuels, also results in a flux of heavy residue. This residue is called petroleum coke or petcoke. At ambient temperature, this fraction is solid and can be ground in traditional mills to obtain a powdered fuel that can be substituted for coal or other traditional fuels. Heavy industries such as cement plants can use these low-cost fuels whose LCV is high at 32,500-35,000 kJ/kg. Over the years, a number of difficulties have come to light in the use of such fuels.

Firstly, petcokes contain low quantities of volatile matters (VM). Consequently, most of the energy will be generated by the oxidation of the carbon residue, also called char, formed during the first step of the combustion process: devolatilisation. When coal is used, typically, 10-60% of the fuel mass is released during devolatilisation under the form of gasses (VM), and approximately, the same fraction of the energy production is due to the gas phase oxidation of the VM. These reactions are much quicker than the oxidation of the carbon residue, also called char. Petcokes contain only 5-16% of VM. Consequently, the residual amount of carbon in the ash can be very high at the exit of a furnace or kiln using petcoke.

Secondly, as is the case for coal, petcokes contain significant amounts of N and S. Their combustion causes the formation of gases containing nitrogen, particularly NO. The relative importance of fuel NO with regard to the thermal NO depends on the rate of fuel nitrogen, on the combustion temperature and on the equivalence ratio in the reactor [START_REF] Glarborg | Some aspects of fuel nitrogen conversion in solid fuel fired systems[END_REF]. In the conditions of the combustion performed in this work, which correspond to those encountered in a fluidised bed reactor or in a cement plant precalciner, only fuel NO is formed [START_REF] Glarborg | Some aspects of fuel nitrogen conversion in solid fuel fired systems[END_REF][START_REF] Thomas | The release of nitrogen oxides during char combustion[END_REF], i.e., NO resulting only from the oxidation of the N initially present in the fuel and not from the oxidation of N 2 present in the air. The sulphur initially present in the petcoke will cause the formation of S-containing gas species, of which SO 2 is the most prevalent [START_REF] Hampartsoumian | Nitrogen sulphur interactions in coal flames[END_REF][START_REF] De Soete | Aspects chimiques de la combustion du charbon pulve ´rise ´-2 e `me partie[END_REF]. The presence of S-species in the gas phase has an effect on the level of NO concentration emitted at the end of the reactor [START_REF] Commandre | Investigating the influence of sulfur on NO formation during the combustion of highly sulphurised petcokes, -6th International conference on technologies and combustion for a clean environment[END_REF][START_REF] Amand | Influence of SO 2 on the NO/N 2 O chemistry in fluidized bed combustion: 1. Full-scale experiments[END_REF][START_REF] Anthony | Relationship between SO 2 and other pollutants emissions from fluidised-bed combustion[END_REF][START_REF] Corley | Postflame behavior of nitrogenous species in the presence of fuel sulfur: II. Rich, CH 4 /He/O 2 flames[END_REF][START_REF] Dam-Johansen | Influence of SO 2 on the NO/N 2 O chemistry in fluidized bed combustion: 2. Interpretation of full-scale observations based on laboratory experiments[END_REF][START_REF] Tullin | Effect of SO 2 and NO on the conversion of fuel nitrogen to N 2 O and NO in single particle combustion of coal[END_REF].

In the past, the use of petcokes in large-scale boilers and furnaces has clearly established that the behaviour of a petcoke can be drastically different from one petcoke to another both in terms of the combustion velocity and of pollutant emissions. Very little literature has been devoted to these fuels, making it difficult to interpret the differences observed in the behaviour of cokes from different sources. The cost of petcokes has fallen to virtually zero, but at the same time, the quality has declined. In particular, sulphur contents have generally risen, and the mineral matter content has also increased.

Developing a tool to predict the behaviour of a petcoke is of course very attractive: this would enable petcokes to be selected that will not produce high amounts of residual carbon in ash and that will keep environmental pollution to a minimum. Inasmuch as these fuels are usually bought in large quantities (several tens of thousands tons of fuel carried by ship), the rewards for developing such a product may be considerable.

This work, carried out over a period of more than 6 years, was devoted to this purpose. As a first step, we characterised a large number of petcokes coming from different origins in terms of all standard laboratory analysis: ultimate analysis, proximate analysis, structural properties, LCV, water content. We then performed combustion of the petcokes using an Entrained Flow Reactor, sometimes also called a Drop tube furnace, which we developed specifically for this work. The temperature was 900 8C, and the gas residence time was 2.13 s. The gas emissions resulting from petcoke combustion were quantified at the exit of the combustion process. We then systematically searched for any correlation that could be established between the properties of the petcoke and its behaviour during combustion. The combustion velocity was quantified through the amount of oxygen that was consumed. The gas species NO and SO 2 were also quantified. The emissions of CO, CO 2 and hydrocarbons (HC) were analysed to see if information about their concentration would facilitate the understanding of the combustion and pollutant formation process. All the details of the procedure developed and the results obtained in this first step are described in this paper.

In parallel, but here termed as a second step, important work was carried out to identify and to characterise the physicochemical phenomena involved in the combustion process and in gas phase pollutant formation. This work, which has already been published [START_REF] Commandre | Formation des oxydes d'azote lors de la combustion de cokes de pe ´trole dans des conditions de pre ´calcinateur de cimenterie[END_REF][START_REF] Salvador | The catalytic effect of vanadium on the reactivity of petcokes with O 2 and NO, 7th International conference on energy for a clean environment[END_REF][START_REF] Salvador | The catalytic effect of vanadium on the reactivity of petroleum cokes with NO[END_REF], will be recalled here and will provide a number of key elements in interpreting and explaining results obtained in the search for correlation between the properties of the petcokes and their behaviour during combustion.

Properties of the petcokes

Twenty-two petroleum cokes were used in this study. They were sampled at different industrial sites, dried, ground in the laboratory and pneumatically sieved to select the 30-100 Am fraction for each fuel. This calibration meant that the particle size distribution for all the petcokes that were burned was very similar. As a result, this parameter will not affect combustion behaviour in our experiments.

Ultimate and proximate analysis was carried out, and results are given in Table 1. The BET surface area and the LCV were also quantified. The 22 petcokes exhibit very similar compositions in terms of the major species: the content in C ranges between 81% and 89%. The volatile matter content ranges between 5% and 16%, which is low when compared with coals, which typically exhibit percentages of 10-60%. The ash contents are also low, inasmuch as all the petcokes contain less than 6% of ash (most of the time around 1%), one of them containing 0.12%.

The amounts of N are comparable with those of coal; they range between 1.21% and 3.5%, most of the values being lower than 2%. It has been demonstrated in previous works [START_REF] Hampartsoumian | Nitrogen sulphur interactions in coal flames[END_REF][START_REF] Commandre | The formation and reduction of NO during the combustion of petroleum cokes: the case of cement plant precalciner conditions[END_REF] that the NO formed during the combustion of powdered fuels in fluidised bed conditions is fuel NO, i.e., NO formed by the oxidation of N in the fuel and not by the oxidation of N 2 present in the air. If all the N of the fuels studied here were converted into NO, the concentration in the flue gas would range between 2317 and 6613 ppm at 3% O 2 .

The sulphur content of the petcokes is significantly higher than that usually present in coal. The values range between 2.5% and 7.5% of S. A full conversion of S into SO 2 , which is known as the very dominant sulphur species after combustion, would range from 2007 to 6400 ppm at 3% O 2 for the present petcoke.

In terms of the heat released by the combustion of the petcoke, the values of the HCV indicated in Table 1 vary by 8.2% around an average value of 34,762 kJ/kg. The reactive surface is known to be an important parameter in the combustion of porous solid fuels. Most of the models used today to describe the combustion of a char residue, representing here 82.0-94.3% of the initial mass, assume that the mass loss rate is proportional to that surface [START_REF] Aarna | A review of the kinetics of the nitric oxide-carbon reaction[END_REF]. Although the reactive surfaces are not always measured in a consistent manner [START_REF] Aarna | A review of the kinetics of the nitric oxide-carbon reaction[END_REF][START_REF] Commandre | The high temperature reaction of carbon with nitric oxide[END_REF], one of the standard ways to estimate this surface is by BET measurement. We used this technique to quantify the specific surface of the char residues after soak pyrolysis, with N 2 as the absorbed gas at 77 K; the results are reported in Table 1. The specific surfaces exhibited are very small; they range between 3.95 and 8.40 m /g, when coals typically present surfaces between 10 and several hundreds of m 2 /g [START_REF] Aarna | A review of the kinetics of the nitric oxide-carbon reaction[END_REF][START_REF] Johnsson | Formation and reduction of nitrogen oxides in fluidized-bed combustion[END_REF].

Experimental device

We carried out the combustion of the 22 petcokes in a device specifically set up to achieve the combustion of pulverised solid fuels under bflashQ conditions, typical of those in a fluidised bed or in a cement plant precalciner. The apparatus consisted of an Entrained Flow Reactor (EFR) whose design details have been described before [START_REF] Van De Steene | Controlling powdered fuel combustion at low temperature in a new entrained flow reactor: laboratory experiments and modelling[END_REF]. The apparatus is shown in Fig. 1. It is an 80-mm i.d. 1-m long isothermal quartz reactor fed with an electrically preheated laminar atmosphere gas flow. The powdered solid fuel is injected through a water-cooled feeding probe and dispersed on the cross-section of the reactor. It burns for a controlled residence time (typically several seconds) before being sampled by a water-cooled nitrogen-quenched probe. Aspecific device was set up to continuously quantify CO, CO 2 , NO X (NO+ NO 2 ) and SO 2 through nondispersive IR absorption and O 2 through paramagnetic analysis.

During our experiments, the atmosphere gas and the reactor walls were heated to 900 8C. Measurements of the gas phase temperature inside the reactor during the combustion of petcoke indicated higher temperatures in the first third of the reactor but not exceeding 930 8C. The atmosphere gas was air, fed at a flow rate of 18 l/min at STP. The petcoke flow rate was 1.5 g/min; in these conditions, the equivalence ratio was 0.8. The sampling probe was placed 900 mm downstream of the feeding probe; the resulting gas residence time was 2.13 s. The difference between these experiments and the conditions found in an industrial boiler or furnace is that the flow was laminar in the EFR when it is usually turbulent on a large scale. Nevertheless, the temperature levels and chemical composition of the flow where similar in the EFR to those in an industrial application.

Results and discussion

The solid fuel was injected continuously until a stable regime was reached; this was indicated by stable values for the gas species concentrations at the sampling point. The results obtained are reported in Table 2. In this table, the O 2 concentration has been ), CO, CO 2 , SO 2 , NO, methanic and nonmethanic hydrocarbons (HC m and HC nm ), total hydrocarbons (HC t ) measured after combustion of petcokes in the EFR (900 8C, 2.13 s, equivalence ratio (0.8);

(1)-gases concentrations converted at 3% O 2 , in daf (dry ash-free basis) O 2cons (%) daf CO (ppm) (1) CO 2 (%) (1) SO 2 (ppm) (1) NO (ppm) (1) HC nm (ppm) (1) HC m (ppm) (1) HC t (ppm) (1) Coke expressed as the difference between the initial O 2 content of the air (20.8%) and the percentage of O 2 at the exit of the reactor so as to deduce the quantity consumed by the combustion. The values ranged between 11.4% and 13.5%; this is equivalent to an average burnout of 80% for all the petcokes (from an O 2 balance).

The first objective of this work was to establish links between petcoke properties and combustion velocity, deduced from the burnout or from the oxygen consumed (O 2cons ). The correlation between O 2cons and each of the properties of the petcokes were systematically investigated by testing linear fitting. In no case was the correlation coefficient r 2 higher than 0.309.

As explained before, one could have expected that the oxygen consumed would have correlated to the BET surface; in fact, as indicated in Fig. 2, r 2 is only 0.037. The reason for this is that the petcokes exhibit very similar BET surfaces, making it unlikely that this parameter would explain the differences in the combustion velocity of the petcokes.

The volatile matter content might also be considered as a parameter to influence the burnout, inasmuch as devolatilisation and oxidation in the volatile matter reactions are much quicker than char oxidation. The plot in Fig. 3 does not, however, show a significant correlation (r 2 =0.011).

The second aim of this work was to establish links between the petcoke properties and pollutant gas emissions. It should be mentioned at this point that all the gas molar fractions specified in Table 2 (except O 2cons ) have been converted at 3% O 2 . This conversion enables concentrations of the gas species that have been formed by the different petcokes with different burnout to be compared, avoiding the dilution effect by unreacted air.

As far as SO 2 emissions are concerned [START_REF] Commandre | Investigating the influence of sulfur on NO formation during the combustion of highly sulphurised petcokes, -6th International conference on technologies and combustion for a clean environment[END_REF], the plot in Fig. 4, where SO 2 emissions are given in relation to the initial content in S of the petcokes, indicates a strong correlation: r 2 =0.96. Moreover, it is possible to establish that the quantity of SO 2 that has been formed is proportional to the fuel burnout. A full conversion of S into SO 2 would theoretically lead to the SO 2 concentrations indicated by the dotted line in Fig. 4. The values obtained Fig. 2. Oxygen consumed during the combustion versus the specific surface of the char residues of the petcokes.

experimentally for a burnout of approximately 80% appear to represent 80% of the full conversion. This observation was reported after investigations into the fate of S in coals [START_REF] Hampartsoumian | Nitrogen sulphur interactions in coal flames[END_REF].

The very high correlation coefficient obtained here also attests to the accuracy of the experiments. This point is particularly important in the context of experiments in the combustion of powdered solid fuels dispersed in a gas phase, which can be perturbed by phenomena such as the fuel-feeding rate fluctuations or bad dispersion of the fuel particles on the cross-section of the reactor.

None of the other fuel properties indicated a correlation to SO 2 emissions during the combustion of the petcokes. The emissions in NO were also confronted with each property of the petcokes with a view of identifying a correlation. The r 2 was never higher than 0.15 (correlation between NO and VM), which indicates that it is not possible to predict NO emissions in a simple way.

The plot in Fig. 5 clearly shows that the initial N content of the petcokes does not determine ultimate NO emissions. This has been reported before in the case of coal [START_REF] Hampartsoumian | Nitrogen sulphur interactions in coal flames[END_REF]. Basically, the fuel nitrogen is released partly in gas phase during devolatilisation and during char oxidation; the NO present in the gas phase can also react with the solid residue to be reduced into N 2 . The kinetic of these reactions varies from one coke to another. The complexity of the phenomena involved explains why NO cannot be predicted only from the knowledge of the initial N content in the fuel. The literature on the combustion of coal and other solid fuels points out that the presence of S species in the gas phase has an effect on the level of NO emissions resulting from the combustion [START_REF] Commandre | Investigating the influence of sulfur on NO formation during the combustion of highly sulphurised petcokes, -6th International conference on technologies and combustion for a clean environment[END_REF][START_REF] Amand | Influence of SO 2 on the NO/N 2 O chemistry in fluidized bed combustion: 1. Full-scale experiments[END_REF][START_REF] Anthony | Relationship between SO 2 and other pollutants emissions from fluidised-bed combustion[END_REF][START_REF] Corley | Postflame behavior of nitrogenous species in the presence of fuel sulfur: II. Rich, CH 4 /He/O 2 flames[END_REF][START_REF] Dam-Johansen | Influence of SO 2 on the NO/N 2 O chemistry in fluidized bed combustion: 2. Interpretation of full-scale observations based on laboratory experiments[END_REF][START_REF] Tullin | Effect of SO 2 and NO on the conversion of fuel nitrogen to N 2 O and NO in single particle combustion of coal[END_REF]. For the case at fuel-lean conditions, which is the case in our experiments, the tendency observed by the authors is that when the sulphur concentration increases, NO concentration decreases. Considering this, we have plotted in Fig. 6 NO emissions versus the S content of the petcokes, which has been shown to be correlated with the SO 2 concentration of the gas phase. The plot does not indicate a correlation in the case of our petcokes, despite the fact that the content in S varies in a range as large as 2.5-7.47%.

As a complement to our search for correlation between the petcoke properties and gas phase pollutant emissions, we investigated the correlation between the different gas emissions themselves. As indicated in Fig. 7, a weak correlation (r 2 =0.43) exists between the combustion burnout, deduced from O 2cons , and NO emissions. It should be remembered that NO concentration is converted at 3% O 2 so as to avoid dilution effects by the air that has not reacted. This relation between the burnout and NO is generally accepted for the fate of N in coals during combustion. Despite the fact that a number of other parameters influence NO emissions, it seems probable from these results that N is released in proportion to the burnout during petcoke combustion.

Catalytic effects

In previously presented work, the approach can be qualified as global in the sense that the gas emissions that were analysed result from a number of reactions that occur simultaneously during the combustion. These reactions are devolatilisation, char residue oxidation, gas phase reactions and also the heterogeneous reaction of a reduction of NO at the porous surface of the char residue. In parallel to the search for correlation, studies have been carried out in which these reactions were considered separately [START_REF] Glarborg | Some aspects of fuel nitrogen conversion in solid fuel fired systems[END_REF][START_REF] Thomas | The release of nitrogen oxides during char combustion[END_REF][START_REF] Commandre | The high temperature reaction of carbon with nitric oxide[END_REF][START_REF] Blair | Evolution of nitrogen and other species during controlled pyrolysis of coal[END_REF][START_REF] Thomas | The release of nitrogen oxides during char combustion[END_REF]. Some of these results are relevant to the present investigation between the petcoke properties and their behaviour during combustion, and are presented here.

The catalytic effect of metals present in a solid fuel is known to be potentially very influential in the combustion process [START_REF] Illan-Gomez | NO reduction by activated carbons: 1. The role of carbon porosity and surface area[END_REF][START_REF] Illan-Gomez | NO reduction by activated carbons: 5. Catalytic effect of iron[END_REF][START_REF] Illan-Gomez | NO reduction by activated carbons: 6. Catalysis by transition metals[END_REF][START_REF] Illan-Gomez | NO reduction by activated carbons: 7. Some mechanistic aspects of uncatalyzed and catalyzed reaction[END_REF][START_REF] Illan-Gomez | Potassium-containing coal chars as catalysts for NO X reduction in the presence of oxygen[END_REF]. We selected 6 petcokes among the 22, covering the range of reactivities observed, and we quantified the following metals: vanadium, copper, molybdenum, iron and manganese. Potassium and calcium were also analysed. The results are grouped in Table 3.

! In Fig. 8, we plotted on the Y-axis the content of all the metals versus O 2cons , the amount of oxygen consumed during the combustion of the petcoke in the EFR. It seems that vanadium, which is present in quantities up to 2570 ppm, has a nonnegligible catalytic effect on the combustion velocity of the petcokes. It was nevertheless found in another work that the combustion of petcokes in a thermogravimetry crucible strongly correlated to the amount of V in the fuel. The reader can refer to Refs. [START_REF] Salvador | The catalytic effect of vanadium on the reactivity of petcokes with O 2 and NO, 7th International conference on energy for a clean environment[END_REF][START_REF] Salvador | Reaction rates for the oxidation of highly sulphurised petroleum cokes: the influence of thermogravimetric conditions and some coke properties[END_REF] for more details. ! The reaction during which NO in the gas phase is reduced at the char surface is known to play an important role in NO balance during the combustion of carbonaceous fuels [START_REF] Aarna | A review of the kinetics of the nitric oxide-carbon reaction[END_REF][START_REF] Johnsson | Formation and reduction of nitrogen oxides in fluidized bed combustion[END_REF][START_REF] Li | Kinetics study of NO and N2O reduction with coal chars[END_REF][START_REF] De Soete | Aspects chimiques de la combustion du charbon pulve ´rise ´-2 e `me partie[END_REF]. Experimental work carried out in the past showed that quantities of 185-849 ppm (starting from 955 ppm at the inlet) could be reduced within several seconds at temperatures of 800-1000 8C [START_REF] Commandre | The high temperature reaction of carbon with nitric oxide[END_REF]. It has been shown previously that the higher the amount of V in the petcoke, the more rapid the reaction is [START_REF] Salvador | The catalytic effect of vanadium on the reactivity of petroleum cokes with NO[END_REF]. This is illustrated in the two figures where we have plotted the amount of vanadium, iron and calcium (Fig. 9) and copper, molybdenum, manganese and potassium (Fig. 10) in relation to the reaction rate for the reduction of NO at the char surface. The reaction rates were calculated from the concentration in NO at the exit of the reactor and were assuming a plug flow reactor [START_REF] Commandre | The high temperature reaction of carbon with nitric oxide[END_REF]. A clear correlation is observed. Consequently, one could expect that the amount of V in the petcoke will determine the amount of NO present in the flue gases of the combustion. Nevertheless, as discussed previously, the NO present in the flue gas of the EFR is the result of a large number of phenomena; this is why the correlation observed here is not explained as a correlation between NO emissions and the amount of V in the petcoke.

Conclusion

A systematic search for correlation between the petcoke properties-in terms of the chemical composition, proximate analysis, physical structure and LCV-and their behaviour during combustion brought the following conclusions:

! The velocity of combustion of a petroleum coke or burnout reached after a given time cannot be correlated with any of the petcoke properties, including the volatile matter content. The catalytic effect of vanadium, which may be present in the petcokes at a mass fraction up to 2570 ppm, has a nonnegligible effect on the burnout. ! The SO 2 emissions are very well correlated (r 2 =0.96) to the initial amount of S in the petcoke. The present work, based upon a large number of petcokes, also confirms a probable release of S from the solid fuel proportionally to particle burnout, as suggested by previous authors. ! The NO emissions cannot be predicted simply from the petcokes properties. They are not correlated with the N content of the fuel. This observation can be understood by considering the large number of phenomena that participate in the formation and reduction of NO and which can be more or less efficient from one petcoke to another. In previous works, we have shown that the reduction of NO at the char surface has an important impact on NO balance, and that this reaction is catalysed by the vanadium entering the composition of the petcoke. The present work shows that this impact is not sufficient to impose a correlation between NO emissions due to the combustion and the content in V of the petcokes.

Fig. 1 .

 1 Fig. 1. Entrained Flow Reactor. (1) Conveyor belt/vibrating corridor/ejector; (2) pneumatic transport; (3) electrical preheater; (4) flow meters; (5) water-cooled feeding probe; (6) dispersion dome; (7) three-zone electrical furnace; (8) 80-mm i.d quartz tube reactor; (9) cyclone collector; (10) exhaust fan; (11) water-cooled nitrogen-quenched sampling probe; (12) particle collector (filter); (13) sampling pump; (14) gas conditioner; (15) gas analysers (infrared/paramagnetic); M-mass flow meters and controllers; N 2 -nitrogen; C.A-compressed Air; W-water (probes cooling).

Fig. 4 .

 4 Fig. 4. Emissions of SO 2 versus the percentage of S in the petcokes.

Fig. 3 .

 3 Fig. 3. Oxygen consumed during the combustion versus the volatile matter content of the petcokes.

Fig. 5 .Fig. 6 .

 56 Fig. 5. Emissions of NO versus the percentage of N in the petcokes.

Fig. 7 .

 7 Fig. 7. Emissions of NO versus the oxygen consumed during the combustion of the petcokes.

Fig. 9 .

 9 Fig. 9. Vanadium, iron and calcium content versus the reaction rate for the reduction of NO at the char surface.

Fig. 10 .

 10 Fig. 10. Metals content versus the reaction rate for the reduction of NO at the char surface.

Table 1

 1 Ultimate analysis, proximate analysis and lower calorific value of the 22 petcokes

		Ultimate analysis				Proximate analysis	LCV	BET
		C %	H %	O %	S %	N %	VM %	Ash %	J/g	m 2 /g
	Coke 1	81.2	3.3		5.97	1.8	11.43	5.88	32,950	3.95
	Coke 2	87.85	3.51	1.1	7.47	1.61	10.48	0.33	34,450	4.5
	Coke 3	87.95	3.71	0.9	7.23	1.51	9.95	0.36	34,400	4.3
	Coke 4	87.9	3.51	1.1	7.08	1.51	10.09	0.38	34,350	10
	Coke 5	87.05	3.89	2.43	4.05	2.03	12.63	1.25	34,500	7.2
	Coke 6	83.96	3.6	1.72	7.14	1.21	10.91	0.91	34,400	6.3
	Coke 7	83.39	3.89	3.11	6.62	1.31	14.47	0.44	33,400	5.5
	Coke 8	87.65	4.16	1.52	4.67	1.52	11.97	1.4	34,900	4.7
	Coke 9	86.69	3.66	1.53	4.37	1.42	9.69	1.75	34,050	8.4
	Coke	82.18	3.75	2.02	5.89	1.32	8.99	1.22	34,300	3.6
	Coke	86.69	3.46	1.88	6.51	1.42	9.41	1.64	34,050	5.3
	Coke	85.5	3.79		4.6	1.46	12.86	1.52	34,300	5.4
	Coke	89.4	3.6		4.02	1.88	10.14	0.47	34,500	5.3
	Coke	88.4	3.7		6.16	1.56	10.96	0.12	34,350	6.2
	Coke	88.5	1.94		6	1.63	5.17	0.5	32,500	6.3
	Coke	88.5	3.6		5.33	1.41	10.01	0.99	34,200	5
	Coke	85.8	3.5		5.85	1.7	11.09	3.35	33,400	4.8
	Coke	87	3.6		5.67	1.7	10.2	0.47	34,250	5.7
	Coke	86	4.2		3.8	1.9	10.37	0.5	34,000	4.7
	Coke	87.4	4.2		4.92	2.4	10.8	0.34	34,400	4.8
	Coke	85.6	3.8		4.96	3.5	16.1	1.9	33,650
	Coke	88.3	3.8		2.5	3.4	14.4	0.5	34,500

Ultimate and proximate analysis are measured on dry basis. The BET specific areas are those of the chars.

Table 2 O

 2 2 consumed (O 2cons

Table 3

 3 Analyses of five metals, K and Ca in six of the petcokes (ppm)

	Sample	V	Co	Mo	Iron	Mn	K	Ca
	2	507	9.1	7.0	615	5.1	45	220
	4	819	-	-	6	-	-	180
	13	2570	1.9	1.4	93	0.5	10	20
	14	1390	3.8	17.3	20	2.6	17	100
	15	899	6.1	26.3	1740	17.1	15	150
	17	1830	3.1	16.6	381	6.1	57	2190

Fig. 8. Metals content in petcokes versus oxygen consumed during the combustion.


H Eslami

Impact of temperature variation on penetration test parameters in compacted soils

Keywords: Laboratory tests, Penetrometer, Temperature, Compacted soils, Illite, Kaolinite-Sand mixture

Impact of temperature variation on

Introduction

In geotechnical engineering the effect of temperature should be clearly studied particularly for applications such as the storage of nuclear waste, the remediation of contaminated sites and the burying of high-voltage cables and thermo-active geostructures. These geostructures are used to provide sustainable heating and cooling by thermal exchange between buildings and the ground through liquids that flow through closed-loop circulation systems integrated into the geostructures. These systems use the subsurface soil as a heating or cooling source for the buildings. Geostructures may be piles, diaphragm walls, tunnel linings, basement slabs or walls [START_REF] Brandl | Energy foundations and other thermo-active ground structures[END_REF][START_REF] Fromentin | Pieux échangeurs: conception et règles de pré-dimensionnement[END_REF][START_REF] Laloui | Comportement d'un pieu bi-fonction, fondation et échangeur de chaleur[END_REF]. The classical way in which thermo-active foundations are used results in cyclic changes in the surrounding ground temperature (approximately 12 °C) over a range of 4 to 30 °C [START_REF] Péron | A method for the geotechnical design of heat exchanger piles[END_REF]. To study the temperature effect on ground structures, the complex interactions between temperature variations and induced stresses or deformations, which may affect long-term structure performance, must be considered [START_REF] Mccartney | Application of Geoexchange Experience to Geothermal Foundations[END_REF].

The thermo-hydro-mechanical behaviour of saturated soils has been studied over the past 45 years, including the effect of temperature on shear strength, yielding and critical-state soil properties. The impact of temperature on the shear strength of soil is complex. A number of studies [START_REF] De Bruyn | The influence of temperature on mechanical characteristics of Boom clay : The results of an initial laboratory programme[END_REF][START_REF] Lingnau | Effects of temperature on strength and compressibility of sand-bentonite buffer[END_REF][START_REF] Mitchell | Shearing resistance of soils as a rate process[END_REF][START_REF] Sherif | Temperature effects on the unconfined shear strength of saturated, cohesive soil[END_REF] have reported that heating decreases the shear strength of soil, whereas others (Abuelnaga, Bergado, & Chaiprakaikeow, 2006;[START_REF] Cekerevac | Experimental study of thermal effects on the mechanical behaviour of a clay[END_REF][START_REF] Houston | Thermo-Mechanical Behavior of Seafloor Sediments[END_REF][START_REF] Laloui | Non-isothermal plasticity model for cyclic behaviour of soils[END_REF][START_REF] Noble | Effect of temperature on strength behavior of cohesive soil[END_REF][START_REF] Tanaka | Stress-strain behaviour of reconstituted illitic clay at different temperatures[END_REF] have obtained the opposite results. These studies show that an increase in temperature has complex and contradictory effects on the shear strength of a soil; for example, a decrease in the viscosity of the pore water due to changes in temperature would tend to decrease the shear strength, whereas the thermal volume change of the soil's components would tend to increase the shear strength. The dominance of one phenomenon or the other depends on the nature of the soil, its initial state (its water content and density) and the applied mechanical load [START_REF] Tang | Effet de la température sur le comportement des barrières de confinement[END_REF][START_REF] Eslami | Impact of the thermo-active geostructures on the hydro-mechanical behavior of the soils[END_REF]. The stress history (over-consolidated or normally consolidated conditions) also has a major impact on the permanent volume change after drainage during heating which modifies the shear strength (e. g. [START_REF] Baldi | Thermal volume changes of the mineralwater system in low-porosity clay soils[END_REF].

In unsaturated soils especially in compacted soils, the number of studies on the effect of temperature variation is limited despite their wide application in the field of geo-environmental engineering. Indeed, studying unsaturated soils is more complex than studying saturated soils, as more parameters need to be considered, including the measurement and control of negative pore water pressure (suction) and temperature and boundary conditions. In compacted soils, both the initial water content and the initial density influence the mechanical behaviour of soils. In fact, these soils do not have the same structure on the dry side and on the wet side of the optimum [START_REF] Lambe | The structure of compacted clay[END_REF].

Therefore, the study of the mechanical behaviour of compacted soils should take into account the initial position of the soil on the compaction curve.

As the properties of soil-structure interface can influence highly the friction and the bearing capacity of the structures, the evaluation of the effect of temperature on the mechanical properties of the soil-structure interface requires also special attention. In the literature, there are various experimental methods under isothermal conditions for studying the behaviour of the soil-structure interface, such as the direct shear test [START_REF] Boulon | Basic features of soil structure interface behaviour[END_REF][START_REF] Desai | Cyclic Testing and Modeling of Interfaces[END_REF] and the simple shear test [START_REF] Kishida | Tests of the interface between sand and steel in the simple shear apparatus[END_REF][START_REF] Zeghal | Soil structure interaction analysis: modeling the interface[END_REF]. But, it is hard to find studies focused on the temperature effect on the soil-structure interface parameters. For example, [START_REF] Stesky | Rock Friction-Effect of Confining Pressure , Temperature , and Pore Pressure[END_REF] reviewed the effect of high temperatures on the friction strength of faulted and jointed rocks. Under some conditions, the friction strength increased with increasing temperature and under other conditions, it remained unchanged or decreased with increasing temperature.

Penetrometer tests involve pushing a probe into the soil and measuring the evolution of cone resistance (q c ) and friction sleeve resistance (f s ). In this study, a minipenetrometer method was developed to quantify evolution of these parameters for two unsaturated compacted soils in response to soil temperature changes.

In the following sections, the mini-penetration test and associated parameters are described. The results of several tests on two different clayey soil samples are presented and discussed. Each one of these soils was compacted at preselected initial dry density and initial water content levels and then submitted to monotonic temperature variations from 1 to 70 °C.

Materials and methods

Different soil samples compacted directly in CBR moulds were tested by measuring the force needed to introduce a probe at a constant penetration rate.

Basic characterisation of materials

Two different types of compacted materials were studied. The first was an illitic soil named Arginotech® from the east of Germany. It contains 77% illite, 10% kaolinite, 12% calcite and traces of quartz and feldspar. Illite is a non-swelling clay present in variable proportions in soils due to mineral illitisation [START_REF] Lynch | Frio Shale Mineralogy and the Stoichiometry of the Smectite-to-Illite Reaction: The Most Important Reaction in Clastic Sedimentary Diagenesis[END_REF]. It has been reported to be sensitive to temperature [START_REF] Tanaka | Stress-strain behaviour of reconstituted illitic clay at different temperatures[END_REF], with an observed decrease in the yield locus of illite with increasing temperature (from 28 to 65 and 100 °C).

The second tested material was a mixture of 50% kaolinite B24 and 50% Hostun sand. Kaolinite B24 is a type of sedimentary clay from the Charente basin in southwest France. Hostun sand is a siliceous sand from the southeast of France. These two materials were oven-dried separately for at least 24 hours (60°C) and then mixed in the desired proportions.

The basic characterisation of each material included determination of its Atterberg limits (AFNOR, 1993) and its Proctor curve (AFNOR, 1999). The liquid limit (LL), the plastic limit (PL), the plasticity index (PI), the optimum water content and the maximum dry density (Proctor compaction) are listed in Table 1 for each material. The optimum water content and the maximum dry density were 31.3% and 1.43 Mg/m 3 , respectively, for the illitic material and 13% and 1.90 Mg/m 3 , respectively, for the kaolinite-sand mixture (Figure 1). Therefore the compaction curves showed that the dry density varies from 1.22 to 1.43 Mg/m 3 for the illitic soil, and from 1.60 to 1.90 Mg/m 3 , for the kaolinite-sand mixture, consequently the mixture has a denser structure.

The particle size distributions of the soils were determined using a laser diffraction particle size analyser (Malvern Mastersizer 2000®) (Figure 2). Almost 85% of the particles of the illitic material are smaller than 0.002 mm (clay particles), and the remaining 15% are smaller than 0.02 mm (silt particles). The particle size distribution of the kaolinite-sand mixture reflects a coarser composition, with 50% clay and silt particles and 50% sand. According to the French standard for soil classification [START_REF] Gtr | Réalisation des remblais et des couches de forme[END_REF], the illitic material is classified as an A3, and the kaolinite-sand mixture is classified as an A1. According to the Unified Soil Classification System [START_REF] Standard | D2487 Standard practice for classification of soils for engineering purposes[END_REF], the illitic material and the kaolinite-sand mixture are classified as a fat clay, CH, and a lean clay with sand, CL, respectively.

Material preparation and compaction

Samples of the two materials were mixed with water to reach the desired water contents and then packed in hermetic bags and left to homogenise for at least 24 hours. The samples with 117 mm in height were compacted in 6 layers in a CBR mould with 152 mm in diameter to ensure the homogeneity of the compacted specimens. Dynamic compaction was applied to each layer. The compaction level was equivalent to the standard Proctor compaction energy. The compaction curve obtained by this specific protocol was superimposed on the standard Proctor curve (Figure 1). The initial temperature of the samples was 20 °C. The samples were thermally cured at 1, 20, 30, 40, 50 and 70 °C for at least 24 h in a climatic chamber while they were packed in hermetic bags so that their water contents were kept constant.

Experimental device

The mini-penetration test is a laboratory test adapted from the standard static penetration test (AFNOR, 1996;[START_REF] Standard | D5778-95 Standard test method for performing electronic friction cone and piezocone penetration testing of soils[END_REF]. A probe consisting of a cone with a 60° apex penetrates the sample at a constant rate of 30 mm/min using a tension-compression system with an accuracy of ± 1N (Figure 3). Two stainless steel probes, termed P1 and P2, were used. The lengths (250 mm), apex angles (60°) and maximal external diameters (12 mm) of the two probes were the same. However, the diameter of the rod of P1 was 10 mm and that of P2 was 12 mm (Figure 3). The maximum penetration length of both probes was 100 mm. The cone resistance was measured with P1, whereas P2 provided a measure of the combined effects of the cone resistance and the friction sleeve resistance.

Three mini-penetration tests were carried out successively in each mould after the temperature equilibrium was reached. All the tests were performed at a temperature controlled room (20±1 °C). The duration of the test was short enough to limit the exchange of temperature during the test. Table 2 shows for example the temperature variation rate during different tests for the illitic material. The position of each penetration test was determined to prevent edge effects (Figure 4). The tests were performed 38 mm from the edge of the mould, and the distance between the test points was 66 mm. The dimensions of the mould were chosen to be large enough to avoid the edge effects, and the diameter of the probes was chosen to be large enough to measure the force. Two successive tests performed with probe P1 in the mould showed the perfect repeatability of the obtained results (Figure 5).

Results

Three different tests were carried out: the penetration of P1, the penetration of P2 and the pull-out of P2. Then, the cone resistance (q c ) as well as the friction sleeve resistance (f s ) during penetration and pull-out were calculated. The coupled influence of initial water content and initial dry density and the influence of temperature on the test results were studied. This part of the test was similar with both probes, as their apex angle and diameter were identical. Part BC is specific to P1. As the diameter of P1 was less than that of its apex, the friction sleeve resistance could be considered negligible from point B onward, and the force reached a constant value Q c over BC for the homogeneous soil. The cone resistance (q c ) was calculated according to equation (1).

Mini-penetration curve typology

q c = Q c / (πR c ²) (1)
where R c is the maximum radius of the conical apex and Q c is the force on the cone.

Part BD of the curve (Figure 6) is specific to the P2 test, in which the measured force (Q t ) is the sum of the force on the cone and the force on the friction sleeve. Q t increased as the probe came into contact with the soil. The friction sleeve resistance (f s ) is the quotient of the difference between Q t and Q c and the lateral surface of the probe in contact with the sample for penetration tests (Equation 2).

f s = (Q t -Q c )/(2πR c (L-L c )) (2)
where L is the penetration length and L c is the height of the cone apex.

After the penetration test, the pull-out test was performed with probe P2. In this case, there was no force on the cone, and the force on the friction sleeve in tension was measured directly. The first part of the curve (EF) shows the high force required to initiate the pull-out. Part FG corresponds to the force on the friction sleeve of the soilprobe interface (Q s * ). Lastly, for part GA, the probe was out of the soil, and there was no friction sleeve resistance. The friction sleeve resistance in tension (f s * ) is the quotient of Q s * and the lateral surface of the probe in contact with the sample for pull-out tests (Equation 3):

f s * = (Q s * )/(2πR c (L-Lc)) (3)

Coupled influence of water content and dry density at 20 °C

In this section, the effects of initial water content and dry density on the measured and calculated strengths at 20 °C are presented. First, the cone resistance results are presented, and then the friction sleeve resistance results for the penetration and pull-out tests are discussed. Table 3 to Table 5 show the water contents (w) and dry densities (ρ d )

of the compacted samples studied in this section. The compaction curves are also presented in the figures to simplify the study of coupled influence of initial water content and dry density.

Cone resistance

Figure 7 (a) presents the P1 penetration force measured for illitic samples prepared at various water contents and dry densities. The samples S4 and S7 were prepared at the same initial dry density (1.38 Mg/m 3 ); therefore, the difference of 190 N between the measured force (320 N for S4 and 130 N for S7) could only be due to the water content difference: w = 29.8% for S4 and w = 32.7% for S7.

The BC parts of the curves are stable, which indicates that the prepared samples are homogeneous. Measured force (Q c ) increased with decreasing water content. Cone resistance (q c ) was then calculated using equation 1(Figure 7 (b)). Then, the mean of the cone resistance obtained in the linear part (BC) for a penetration depth between 50 mm and 100 mm in the P1 test was used as the mean cone resistance for the test. These limits were selected because the values of cone resistance are homogenous and stable.

Figure 8 superimposes the mean cone resistance (q c ) onto the compaction curve for each material.

 For the illitic samples, the maximum values were obtained for the materials on the dry side (q c = 11,270 kPa), and q c decreased linearly when approaching the optimum

(1,800 kPa). On the wet side of the optimum, q c changed slowly, trending toward approximately 1,000 kPa (Figure 8(a)).

 For the kaolinite-sand mixture, a rapid decrease in q c on the dry side of the optimum and a slow decrease on the wet side were observed. The maximum value was also located on the dry side (q c = 7,055 kPa) (Figure 8(b)).

For the kaolinite-sand mixture, the cone resistance at the optimum (3,800 kPa) was higher than for the illitic material (1,800 kPa). This could be due to different nature of the studied soils and the high density of the mixture (1.90 Mg/m 3 for the kaolinitesand mixture and 1.43 Mg/m 3 for the illitic material). However, the behaviour of the two studied materials was similar overall, as the maximum values of q c were always on the dry side of the optimum water content, and there was a pronounced decrease in the measured parameters on the dry side of the optimum and a slow decrease on the wet side of the optimum.

Friction sleeve resistance for penetration tests

Figure 9 (a) presents the total force (Q t ) using probe P2 (curve ABD). Q t decreased with increasing water content. This result is coherent with the evolution of the force on the cone (Q c ). The difference between the curves obtained with P1 (Figure 7 (a)) and the curves obtained with P2 (Figure 9 (a)) provided the force on the friction sleeve (Figure 9 (b)). f s increased with penetration depth due to an increase in the soil-probe interface. For each test and each penetration depth, friction sleeve resistance (f s ) was calculated (Equation 2) (Figure 10). The mean values of f s for penetration depths between 50 mm to 100 mm were used to obtain the mean friction sleeve resistance for the test (Figure 11).

 For the illitic material, there was a peak in the friction sleeve resistance according to the soil density and water content evolution. This could be due to the compacted soil microstructure, which is irregularly aggregated on the dry side [START_REF] Lambe | The structure of compacted clay[END_REF]. In fact, on the dry side and for low water contents, the effective contact surface between the probe and the soil is small and irregular. As the initial water content increases, the contact surface increases, and the friction sleeve resistance also increases until the peak. From the peak value, fs decreases because of the lubrication due to the increase in water content.

For the illitic material, a maximum of 420 kPa for fs was reached on the dry side of the optimum for the sample prepared with a 27.9% water content (3.4% less than the optimum water content). After that point, f s decreased quickly, reaching 76 kPa for samples prepared with a 32.7% water content (optimum water content + 1.4%). On the wet side of the optimum, the decrease in f s was smaller than that on the dry side (Figure 11Figure 11(a)).

 For the kaolinite-sand mixture, the maximum measured friction sleeve resistance was 180 kPa; this was less than that of the illitic material, which could be due to the nature of the material. As with the illitic material, friction sleeve resistance decreased quickly on the dry side but slowly on the wet side, and it reached a minimum of 13 kPa (Figure 11(b)). Comparing the friction sleeve resistances at different water contents and dry densities yields the same conclusion as the examination of the cone resistance results for the kaolinite-sand mixture samples: as the water content decreases, the friction sleeve resistance increases.

The increase in the studied parameters (cone resistance and friction sleeve resistance) with decreasing water content on the dry side of the optimum water content could be related to the increase in the matric suction, which represents the combined effects of the forces holding the water in the soil.

Given the correlation between shear strength (modulus) and penetration test parameters, the results can be compared with those in the literature. The observed trends for the studied parameters in the present study are consistent with the results obtained by [START_REF] Sawangsuriya | Modulus-suction-moisture relationship for compacted soils[END_REF] for the normalised small-strain shear modulus of certain soils. The soils tested were a lean clay, a silt and a clayey sand, and they were compacted over a range of moulding water contents from 4% dry of optimum to 4% wet of optimum. They determined the modulus based on a shear wave propagation technique using bender elements. They observed that the shear modulus increases on the dry side of optimum with increasing suction and decreasing moulding water content. They also found that if any peak occurred, it was on the dry side of the optimum. These results are also consistent with those obtained by [START_REF] Wiebe | Influence of pressure, saturation, and temperature on the behaviour of unsaturated sand-bentonite[END_REF], who studied the evolution of the shear strength for a sand-bentonite mixture prepared at various degrees of saturation (S r ). Tests carried out at the same temperature showed that as S r decreased from 98% to 50%, shear strength increased. However, for a decrease in S r from 50% to 35%, they recorded a decrease in shear strength. They attributed this behaviour to the variation in suction and the deformability. In fact, at high water contents (S r = 98%), suctions are low and aggregates are deformable, which results in low shear strengths. At low water contents (S r = 50%), the suction is high, and therefore, the increase in the internal stiffness leads to a higher shear strength. For the lowest water contents (S r = 35%), the decrease in the shear strength is due to the decrease in the stiffness of aggregates, which become more brittle. [START_REF] Uchaipichat | Experimental investigation of thermo-hydromechanical behaviour of an unsaturated silt[END_REF] also studied the shear strength of a silt under different suctions (0, 100 and 300 kPa). The deviatoric stress increased with increasing suction, and the maximum deviator stress occurred in a shear test with an initial matric suction of 300 kPa. They concluded that the expansion of the yield locus with increasing suction could be due to the suction-hardening phenomenon in unsaturated soils.

Friction sleeve resistance for pull-out tests

Figure 12(a) illustrates the force increase with decreasing water content in the pull-out tests. These results are similar to those obtained for the penetration tests. As with the penetration tests, the measured forces for the mixture are lower than those of the illitic material. For each test and each penetration depth, the friction sleeve resistance for the pull-out test (f s * ) was calculated using Equation 3 (Figure 12Figure 12(b)). The mean value of f s * at penetration depths of 80 mm to 90 mm was taken as the friction sleeve resistance for the pull-out test (Figure 13) because, for this parameter, the values were constant and homogeneous.

 For the illitic material, as with the penetration test, there was a peak for the friction sleeve for the same sample, with a maximum absolute value of friction sleeve resistance of 280 kPa on the dry side of the optimum. The absolute values of friction sleeve resistance (f s * ) decreased, reaching 52 kPa for samples prepared on the wet side of the optimum. (Figure 13(a)). The existence of a peak for friction sleeve resistance for the illitic material could be due to the evolution of the force on the friction sleeve, as explained for the penetration tests in the previous section.

 For the kaolinite-sand mixture, the absolute values of the evolution of friction sleeve resistance are the same as in the penetration test, starting from a maximum absolute value of friction sleeve resistance (140 kPa) and reaching the small values on the wet side of the optimum (Figure 13(b)).

Coupled influence of water content and dry density at different temperatures

Some samples of illitic material and kaolinite-sand mixture were prepared at various water content and dry density then cooled down to 1°C or heated to 30, 40, 50 or 70°C

to study the effect of monotonic temperature variation (Table 4 andTable 5).

Penetration and pull-out tests were performed and cone resistance and friction sleeve resistance were calculated for each sample as explained in the section 3.2. In this section the final results are presented and discussed.

 For the illitic material, 12 tests at 1 °C, 11 tests at 40 °C and 4 tests at 70 °C were performed. Figure 14(a) presents the mean cone resistance (q c ) superimposed onto the compaction curve at different temperatures. As with the tests at 20 °C, the maximum values of the cone resistance were obtained on the dry side of the optimum for all temperature values. The peak values of the friction sleeve resistance were obtained on the dry side of the optimum in the penetration and pull-out tests (Figure 15(a) and

Figure 16(a)). For the tests at 1 °C, all the studied parameters were close to those obtained at 20 °C, therefore it is hard to deduce a potential impact of temperature. For the tests at 40°C, cone resistance and friction sleeve resistance were not affected by temperature variations for the samples prepared on the wet side of the optimum. In contrast, for the samples prepared on the dry side of the optimum, increasing temperature tended to decrease friction sleeve resistance. Cone resistance changed slightly with increasing temperature but did not vary significantly (Figure 14(a), Figure 15(a) and Figure 16(a)). The tests at 70 °C carried out on the dry side of the optimum were similar to the obtained results at 40 °C: as the temperature increased, the studied parameters decreased. This temperature impact was more evident for friction sleeve resistance than for the cone resistance. These results are consistent with those obtained by [START_REF] Wiebe | Influence of pressure, saturation, and temperature on the behaviour of unsaturated sand-bentonite[END_REF], who studied the effect of temperature on soil shear strength.

Their results showed that shear strength decreases with increasing temperature and that the effect of temperature is more marked when the degree of saturation is low, i.e., on the dry side of the optimum. They concluded that the decrease in resistance is caused by the decrease in suction due to heating. The results of [START_REF] Uchaipichat | Experimental investigation of thermo-hydromechanical behaviour of an unsaturated silt[END_REF] temperature and suction controlled shear tests showed a decrease in shear stress with increasing temperature for the same suction. Increasing temperature caused a decrease in the yield locus due to the thermal softening. The decrease in the studied parameters with increasing temperature could be due to the reorganisation of the soil microstructure and the decreased viscosity of water. Indeed, the decrease in pore water viscosity can facilitate sliding between grains and the soil-probe interface, thereby reducing the measured strengths. This could explain the observed temperature effect on the dry side of optimum for the compacted illitic material.

 For the kaolinite-sand mixture, 3 tests at 1 °C, 3 tests at 30 °C and 2 tests at 50 °C were performed. The variation of the parameters according to the soil density and water content evolution is similar to 20 °C at 1, 30 and 50 °C (Figure 14 4

Conclusions

The purpose of this study was to propose a simple method to measure the effects of temperature variations on the cone resistance (q c ) and friction sleeve resistance (f s ) of compacted soil samples prepared at various initial water contents and densities. A minipenetration test was developed, and the results obtained showed that the test is repeatable and reliable. For the studied compacted soils, decreasing the water content increased both cone resistance and friction sleeve resistance. Overall, the dry density of the soil seems to have less impact on the penetration parameters than the water content.

Both cone resistance and friction sleeve resistance vary more sharply on the dry side of the compaction curve than on the wet side.

For the illitic material compacted on the dry side of the optimum water content, a slight decrease in cone resistance and a large decrease in friction sleeve resistance were observed with increasing temperature. On the wet side, the changes in cone resistance and friction sleeve resistance with changes in temperature were not significant. For the kaolinite-sand mixture, on both the wet and dry sides of the compaction curve, the temperature effect on the penetration parameters was negligible.

Indeed, the initial water contents and dry densities, as well as the nature of the soil, could be considered the most important factors affecting the influence of temperature variations on the studied parameters.

Of note, the samples in this study were maintained at constant water contents while the temperature was varied. The coupled evolution of the temperature and the water content should also be considered.

ACKNOWLEDGEMENTS

This

Figure 6

 6 Figure 6 presents an example of obtained curves for different tests using probes P1 and

  (b), Figure 15(b) and Figure 16(b)). Cone resistance was not strongly affected by temperature variations between 1 and 50°C (Figure 14(b)). For the friction sleeve resistance in the penetration and pull-out tests, the evolution of the temperature did not affect the studied parameters, and excluding two mentioned tests for cone resistance, there was no significant variation with temperature (1, 20, 30 and 50 °C), as shown in Figure 15(b) and Figure 16(b). This could be due to the high density of the mixture limiting the reorganisation of the soil microstructure.

Figure 1 :

 1 Figure 1: (a) Specific and standard compaction curves of the kaolinite-sand mixture. (b) Specific compaction curve of the illitic material. Sr: The degree of saturation.

Figure 3 :

 3 Figure 3: Mini-penetration test with the tension-compression machine and the probes for the mini-penetration test.

Figure 6 :

 6 Figure 6: Force versus depth measured using probes P1 and P2 for the penetration test (ABC and ABD) and using probe P2 for the pull-out test (EFGA).

Figure 7 :

 7 Figure 7: (a) Force measured along the penetration of P1 for samples S1 to S7. (b) Evolution of cone resistance for samples S1 to S7.

Figure 8 :

 8 Figure 8: Mean cone resistances (q c ) at different water contents and dry densities for the P1 tests: (a) illitic material and (b) kaolinite-sand mixture.

Figure 9 :

 9 Figure 9: (a) Evolution of the total force (P2 tests) and (b) the deduced force on the friction sleeve (P2-P1) for samples S1 to S7.

Figure 11 :

 11 Figure 11: Friction sleeve resistances (f s ) at different water contents and dry densities: (a) illitic material and (b) kaolinite-sand mixture.

Figure 12 :

 12 Figure 12: (a) Friction forces for samples S1 to S7 for the pull-out tests. (b) Evolution of the friction sleeve resistance (f s * ) for samples S1 to S7 for the pull-out tests.

Figure 13 :

 13 Figure 13: Mean friction sleeve resistances for the pull-out tests (f s *) at different water contents and dry densities: (a) illitic material and (b) kaolinite-sand mixture.

Figure 14 :

 14 Figure 14: Mean cone tip resistances (q c ) at different water contents and dry densities: (a) illitic material at 1, 20, 40 and 70 °C and (b) kaolinite-sand mixture at 1, 20, 30 and 50 °C.

Figure 15 :

 15 Figure 15: Mean friction sleeve resistances (f s ) at different water contents and dry densities: (a) illitic material at 1, 20, 40 and 70 °C and (b) kaolinite-sand mixture at 1, 20, 30 and 50 °C.

Table 1 :

 1 study was part of the GECKO research program funded by ANR. Characteristics of illitic material and kaolinite-sand mixture

	6 BIBLIOGRAPHY
	Abuelnaga, H., Bergado, D., & Chaiprakaikeow, S. (2006). Innovative thermal
	technique for enhancing the performance of prefabricated vertical drain during the
	preloading process. Geotextiles and Geomembranes, 24(6), 359-370.
	AFNOR. (1993). NF P94-051 Sols: reconnaissance et essais ; Détermination des limites
	d'Atterberg -Limite de liquidité à la coupelle-Limite de plasticité au rouleau. (p.
	15). Paris, France: Association Française de Normalisation.
	AFNOR. (1996). NF P94-113 Sols: reconnaissance et essais ; Essai de pénétration
	statique (p. 16). Paris, France: Association Française de Normalisation.

AFNOR.

(1999)

. NF P 94-093 Sols : Reconnaissance et essais Détermination des références de compactage d 'un matériau. Essai Proctor Normal-Essai Proctor Modifié (p. 18). Paris, France: Association Française de Normalisation.

Table 2 :

 2 The initial and final temperature on the centre of the sample of illitic material.

	Initial temperature (°C)	1	20	40	70
	Temperature variation rate (°C/min)	0.17	0	0.20	0.53

Table 3 :

 3 Water contents (w) and dry densities (ρ d ) of illitic samples for

	the results presented in figures 7, 9, 10 and 12.
	Sample	w (%) ρ d (Mg/m 3 ) q c (kPa)	f s (kPa)	f s * (kPa)
	S1	25.8	1.25	7384	368	-
	S2	27.9	1.34	4446	418	-281
	S3	28.9	1.36	3621	328	-220
	S4	29.8	1.38	2808	227	-152
	S5	30.4	1.43	2158	145	-100
	S6	31.9	1.41	1519	88.5	-66
	S7	32.7	1.38	1147	76	-52

Table 4 :

 4 Water contents (w) and dry densities (ρ d ) for illitic material

	samples.
	Temperature (°C) w (%) ρ d (Mg/m 3 ) q c (kPa)	f s (kPa)	f s * (kPa)
		23.6	1.28	11270	380	-252
		25.8	1.25	7384	368	-
		26.1	1.29	7003	403	-229
		26.4	1.27	6871	328	-202
		26.8	1.30	5747	392	-
	20	27.8	1.33	5029	410	-
		27.9	1.34	4446	420	-280
		28.9	1.36	3621	328	-220
		29.8	1.38	2808	227	-152
		30.4	1.43	2158	145	-100
		31.9	1.43	1519	89	-66
		32.7	1.38	1147	76	-52
		24.1	1.24	8158	353	-205
		26.9	1.29	6117	391	-212
		28.1	1.33	4035	366	-247
		28.9	1.37	3015	296	-201
		29.1	1.36	3498	294	-206
	1	30.2 30.5	1.39 1.39	1850 1970	150 138	-99 -96
		31.6	1.42	1041	89	-58
		32.3	1.41	1331	86	-60
		32.3	1.42	1346	77	-56
		32.3	1.39	1158	66	-47
		34.2	1.39	661	37	-29
		25.0	1.26	7377	326	-185
		26.2	1.31	5595	293	-193
		28.1	1.31	3644	313	-196
		28.2	1.33	3710	283	-197
		29.5	1.40	2559	223	-139
	40	29.6	1.39	2480	192	-134
		29.6	1.37	2424	189	-122
		30.6	1.42	2094	134	-93
		31.0	1.43	1935	121	-83
		31.7	1.40	1530	91	-63
		33.0	1.38	1064	64	-44
		21.4	1.22	8250	289	-136
	70	24.5 26.0	1.29 1.31	7524 5300	284 296	-150 -167
		28.0	1.36	3938	319	-215

Table 5 :

 5 Water contents (w) and dry densities (ρ d ) for kaolinite-sand mixture samples.

	Temperature (°C) w (%) ρ d (Mg/m 3 ) q c (kPa)	f s (kPa)	f s * (kPa)
		11.3	1.74	7055	-	-
		11.9	1.82	5637	-	-
		12.3	1.87	4556	180	-140
	20	13.0 14.2	1.88 1.85	3514 1287	141 54	-90 -42
		15.1	1.82	845	30	-24
		16.9	1.73	393	18	-12
		18.5	1.66	198	13	-8
		12.0	1.82	5863	191	-182
	1	13.4	1.90	2320	100	-70
		15.3	1.79	802	31	-26
		11.7	1.81	5564	207	-177
	30	13.2	1.90	2300	93	-66
		15.0	1.79	794	36	-26
	50	12.2 12.5	1.85 1.87	3467 4231	160 155	-105 -100