Energy piles are bi-functional structural elements that are used to support the structural loads of a building and to operate as a geo-heat exchanger for shallow geothermal energy systems. In urban areas, energy piles can eventually be used to harvest the heat loss through the basement enclosure and re-inject the energy to the building for the heating and cooling purposes. In spite of a higher thermal proﬁle beneath the buildings due to energy loss through the below-grade envelope, an underground thermal imbalance may still occur due to the application of energy piles in cold regions. This paper aims to study the structural performance of an energy pile in cold regions considering the potential occurrence of thermal imbalance in the foundation soil through a Thermo-Mechanical (TM) analysis. This will be the ﬁrst step of adoption of such a technology in Canadian cold climate. Results showed a safe margin between the pile settlement and the allowable settlement. However, the axial stress applied to the pile increased by 9% due to thermal loads. In addition, a maximum decrease of 9% in mobilized shaft friction along the pile-soil interface was recorded due to the thermal loads. To consider the energy piles as an alternative for buildings energy supply in cold regions, further considerations should be made to keep the mechanical response of the energy piles in the admissible range. Keywords: Geothermal pile, Thermal imbalance, Geothermal energy, TM model, cold regions 1. Introduction The energy consumption has increased all over the world due to the population growth and industrial needs. One of the alternatives to address this increasing energy demand is the application of new technologies to har- ness the energy stored in the ground due to the moderate Corresponding author Email address: Pooneh.Maghoul@umanitoba.ca (Pooneh Maghoul)temperature of the ground all over the year [1, 2]. One of the most promising technologies for the application of geothermal energy is the Ground Source Heat Pump (GSHP). Such a system consists of a heat pump, ground heat exchangers, and a heat distribution system. The ground heat exchangers (GHEs) can be an open-loop or a closed-loop system and its selection depends on vari- ous parameters that impact the feasibility of the project Preprint submitted to Elsevier November 30, 2020 such as the availability of natural groundwater, ponds or lakes. Despite the simple design and better thermo- dynamic performance of an open-loop system, environ- mental regulations restrict the application of this type of GHE. In fact, changes in groundwater temperature dur- ing the operation of GSHP may result in an increase in the bacterial content and suspended particles in ground- water [3]. On the other hand, closed-loop systems can be in- stalled horizontally (straightly horizontal or spirally horizontal loops placed in trenches) or vertically (U- shaped, W-shaped or spiral loops placed in boreholes) depending on the availability of land and energy de- mand. One of the disadvantages of the application of horizontal loops in cold regions like Canada is that the ground temperature is subjected to seasonal variation at shallow depths [4]. However, the installation costs of vertical closed-loop systems can be fairly highly de- pendent on depth and the required number of boreholes [5]. To overcome this issue, a rapidly growing con- struction technology, which integrates the geothermal energy technique into the structural pile foundation sys- tem has been developed [6–9]. The geothermal pile op- eration principle is based on the circulation of a heat carrier ﬂuid in polyethylene pipes embedded in the piles through which heat exchange occurs between the energy pile and the surrounding medium [10–13]. This energy further feeds the GSHP system so that the energy pile acts as heat exchangers in addition to their main func- tion as structural support elements. Geothermal piles are broadly applied in Europeancountries [9, 14–18]. For example, Bourne-Webb et al. [14] conducted some in-situ tests on a single energy pile installed in London Clay. The main scope of the in- situ tests was to evaluate the mechanical performance of the pile subjected to thermal cycles. The temperature and strain proﬁles of the tested pile were recorded using an optical sensor system, vibrating-wire strain gauges, thermistors, and external load control elements. Based on the obtained results, variations in the temperature of the pile, depending on whether the pile is heated or cooled, changed the mobilized shaft friction along the pile-soil interface. It was speculated that if the margin between the pile ultimate shaft resistance and shaft fric- tion along the soil-pile interface is large enough, the pile capacity will not change signiﬁcantly. Although full-scale in-situ experiments are more re- liable to evaluate the TM response of energy piles, they usually require a fair amount of budget and time. An alternative approach to estimate the behavior of the en- ergy piles is analytical or numerical studies. For exam- ple, Pasten and Santamarina [19] evaluated the mechan- ical behavior of energy piles using the one-dimensional load transfer method. This study emphasized on the im- portance of the ﬁrst few cycles of thermal loading in assessing the pile head displacement and corresponding shaft friction variations. Jeong et al. [20] carried out a three-dimensional (3D) ﬁnite element analysis to quantify the axial load and set- tlement of energy piles in a group which are subjected to TM loading. This study highlighted the importance of soil type on the mechanical response of the energy 2 piles. It was concluded that the head displacement of the energy piles embedded in clay was up to 1.5 mm larger than the head displacement of piles installed in sand although both cases were in acceptable limits. One of the signiﬁcant advantages of numerical anal- ysis over in-situ tests is the potential for studying the behavior of energy piles over long-term periods. Olgun et al. [21] studied the long-term displacement of di er- ent arrangements of energy piles over a 30-year span of time under di erent climatic conditions. Impact of climate condition manifested itself as a variation in the ground temperature as well as a variation of energy de- mand of the building for which the energy foundation was designed. This factor has been studied by Saaly et al. [22] Liu et al. [23], Dacquaya et al. [24], and An- dolﬁ et al. [16]. Although broadly expanded in Europe and the USA, in Canada, the application of geothermal piles is not well established and limited knowledge regarding the thermal and mechanical operation of such systems ex- ists. In Canada, the operation of GSHP systems may not be satisfactory and may degrade over a few years of the operation of energy piles [25]. The reason for such malfunction is underground thermal imbalance, which is the term used for the gradual decrease in the ground temperature over years because of an excessive heat ex- traction for heating the space during the long winter season. As soil temperature decreases, the outlet tem- perature of the GHE lowers gradually. This contributes to the deterioration of the heating performance of the GSHP. In addition, the underground thermal imbalanceand excessive heat extraction can decrease the soil tem- perature below the freezing point. This causes freezing at the pile-soil interface which temporarily increases the bearing capacity of the energy pile. However, during spring, thawing occurs at the soil-pile interface, which decreases the bearing capacity of the piles signiﬁcantly. The occurrence of such freezing and thawing cycles at the soil-pile interface adversely a ects the structural in- tegrity as well as the energy e ciency of such systems [26]. Therefore, it is of paramount importance to con- sider the potential occurrence of ground thermal im- balance, i.e. the energy performance, in the design of geothermal pile foundation systems in cold regions. In urban areas, the ground temperature below the buildings base area is fairly higher in comparison to open ﬁelds due to the heat dissipation of buildings through their sub-grade enclosures [27]. For exam- ple, Saaly et al. [28] showed that for an institutional building in the campus of the University of Manitoba with approximate annual heat loss of 30 MWh through the below-grade envelope, soil temperature could in- crease up to 5C comparing to the situation that no heat loss occurs through the building’s basement. Such heat source can only be targeted by energy piles since the hot spot of each heat loss is beneath the building. Therefore, the lost energy can be re-harvested using a geothermal foundation system which favorably a ects the eciency of the GSHPs. Moreover, this additional dissipated heat from the building basement can ame- liorate the potential thermal imbalance in the ground. In this regard, Saaly et al. [22] conducted a study to 3 determine the thermal capacity of a proposed hypo- thetical geothermal foundation system considering the heat loss through the basement of an institutional build- ing in cold regions. The energy piles are subjected to both thermal loads due to temperature variations of the ﬂuid heat exchanger whose magnitude is a func- tion of the building heating and cooling demand as well as structural loads [29]. Therefore, aside from the en- ergy performance of the system, the impacts of ther- mal loading on the structural performance of geother- mal piles should be assessed. For example, thermal loads cause volumetric expansion and contraction when piles are heated and cooled, respectively. In addition, the thermo-mechanical properties of soils surrounding the pile foundation are temperature-dependent, which can a ect the interaction between the soil and founda- tion at their interface [30]. Although above-mentioned numerical works pro- posed the geothermal system for re-harvest of buildings energy loss, mechanical response of the proposed en- ergy piles used in buildings foundation yet needs to be investigated. Therefore, this paper aims to investigate the thermo-mechanical behavior of a single pile in an energy foundation system, which is designed for energy supply of an institutional building in Winnipeg, Canada. This foundation system is subjected to heat dissipation through the below-grade enclosure of the building. The present work is intended to provide a tool for prediction of the thermo-mechanical response of energy piles by incorporating the knowledge of buildings energy loss.2. Features of the Energy Foundation System The energy foundation system belongs to a four story institutional building, the Stanley Pauley Engineering Building (SPEB), with a total base area of 1,150 m2 located in Winnipeg (MB), Canada. Winnipeg expe- riences an average daily temperature ﬂuctuation that ranges from +20C in summer to -20C in winter [31]. Consequently, the buildings in Winnipeg usually have a high energy demand, both for heating and cooling pur- poses [32]. Winnipeg soil is mainly composed of ﬁne-grained de- posits of Lake Agassiz that are followed by Precambrian basement of bedrock [27]. Based on the site investi- gation reports, the stratigraphy of the site consists of a layer of dark grey clay deposits having a thickness of 2 m overlaying a 14.7 m thick layer of brown clay fol- lowed by a dense layer of till deposits [33]. The foundation of the building consists of 119 con- crete micro-piles designed for an ultimate limit state axial resistance of 1000 kN, which are embedded at the depth of -4.25 m below grade. Saaly et al. [22] proposed a geothermal foundation system in which all the 119 piles were considered to be energy piles and were equipped with heat carrier pipes. In other words, all piles are assumed to operate as a geothermal heat exchanger to supply the SBEB energy demand (Fig. 1). To this end, the foundation piles are equipped with heat carrier pipes to exchange heat with the surrounding medium. To determine the inlet temperature of ﬂuid in the heat 4 Figure 1: Monthly heating and cooling energy demand of the Stanley Pauley Engineering Building [32]. exchanger pipe, two main factors have been considered: (i) meeting the thermal balance of the ground during the heat extraction and rejection, (ii) avoiding freezing at the soil-pile interface. Considering the mentioned criteria, the ﬂow rate of the heat exchanger ﬂuid en- tering the closed loop embedded in the pile is chosen as 30.5mo lit /min while its inlet temperature varies be- tween -1C during Nov–Apr and 20C during May–Oct. Based on the results, up to 41% of the energy demand of the Stanley Pauly building can be supplied using an energy pile foundation system by meeting the ground thermal balance and avoiding the freezing at the soil- pile interface [22]. Knowing the inlet temperature of the heat exchanger ﬂuid, the mechanical behavior of a single geothermal pile in the foundation system of the SPEB is evaluated in the following sections.3. Thermo-Mechanical Model 3.1. Governing Equations A coupled thermo-mechanical model has been ap- plied to assess the mechanical and thermal behaviors of one of the piles of the SPEB foundation equipped with a heat exchanger pipe. The governing equations applied to study the thermo-mechanical behavior of the geother- mal pile in contact with the surrounding soils are based on the following assumptions: Soil is considered to be fully-saturated having two- phases (solid skeleton and water). The soil is mod- eled as an isotropic, elastoplastic material. The Mohr-Coulomb constitutive model is used to de- termine the size of the yield surface in the soil. Pile material is taken to be solid, isotropic and lin- ear elastic. Perfect contact has been assumed at the interface between the soil and pile. The solid grains and ﬂuid phase of the soil are modeled as incompressible materials. To consider the equilibrium of forces in the domain, the conservation of linear momentum in absence of in- ertial terms is applied: r:+b=0 (1) wherer:is the divergence operator, is the total (Cauchy) stress tensor and bis the body force vector which can be determined as shown below: 5 b=g (2) where, gis the gravitational acceleration vector and is the average bulk density of the soil which can be deter- mined by the following equation: =(1n)s+nw (3) in whichsandware the density of solid skeleton and water, respectively. nis the porosity of the soil. The constitutive law of the solid skeleton describing the stress-strain relations is deﬁned in terms of the ef- fective stress as described in Eq. 4. 0=De(""T"pl) (4) where Deis the elastic sti ness tensor, "is the strain tensor,"Tis the thermal strain tensor, and "plis the plastic strain tensor. Plastic strains are computed by the Mohr-Coulomb yield criterion. Also, the thermal strains are computed using the following equation: "T=1 3( srT)I (5) in which"Tis the thermal strain tensor, Iis the second order unit tensor, sis the volumetric thermal expansion coecient,ris the gradient operator, and Tstands for the temperature. To determine the heat transfer in porous media, the energy conservation equation is described below:C@T @t+r:qc=0 (6) whereCis the volumetric heat capacity of the soil which is a function of volumetric heat capacity of solid skeleton (sCs) and pore water ( wCw) as described be- low: C=(1n)sCs+nwCw (7) In Eq.7, qc[W/m2] is the conductive heat ﬂux through the volume which is deﬁned by Fourier’s law as follows: qc=rT (8) wheredenotes the e ective thermal conductivity of the material whose value for the soil can be determined as described in Eq. 9. =(1n)s+nw (9) in whichis the e ective thermal conductivity of the soil,sis thermal conductivity of solid grains and wis thermal conductivity of water. 3.2. Numerical Model The numerical model has been implemented in the ﬁnite-element commercial software COMSOL Multi- physics. The dimension of the 2D axisymmetric model applied in this study is 10 m 10 m. On top of the soil domain, a 0.23 m thick layer of thermal insulation and a 0.2 m thick layer of concrete slab are placed (Fig. 2). The modeled pile is a friction pile which has a diameter of 0.4 m and a length of 5 m. The thermo-mechanical 6 properties of the materials used in the model (soil (in wet condition), concrete and insulation) are summarized in Table 1. Based on the structural and design documents, 5 m micro-piles have been used in the foundation of the SPEB considering the structural load and geotechnical parameters. Figure 2: Axisymmetric ﬁnite elements model. The boundary conditions of the thermo-mechanical (TM) model are as follows: The thermal boundary conditions: Constant tem- perature on the heat exchanger, adiabatic condition at the bottom and surrounding boundaries are con- sidered. Also, to estimate the impacts of heat dis- sipation through the building basement slab, a con- vective heat ﬂux q0[W=m2] is assigned to the top boundary of the model as follows: q0=hTindoorTs(10) where his the convective heat transfer coe cient whose value is assumed to be equal to 6 W /m2°C [26], Tindoor is the basement indoor air temperaturethat is assumed constant and equal to 20°C, Ts[C] is the temperature at the ground surface. The mechanical boundary conditions: The roller condition on the right boundary of the model, and ﬁxed constraint at the bottom of the domain are considered. An axisymmetric condition is assumed from the left boundary. Soil Young’s modulus [Pa] 70106 Poisson ratio [-] 0.236 Density [kg /m3] 1767 Thermal conductivity [W /m°C] 1.483 Speciﬁc heat capacity [J /kg°C] 960 Thermal expansion coe cient [1 /K] 1.7Ö105 Porosity [-] 0.535 Cohesion [kPa] 5 Friction angle [] 18 Concrete Young’s modulus [Pa] 40109 Poisson ratio [-] 0.2 Density [kg /m3] 2300 Thermal conductivity [W /m°C] 1.5 Speciﬁc heat capacity [J /kg°C] 880 Thermal expansion coe cient [1 /K] 8.5Ö106 Insulation Density [kg /m3] 34 Thermal conductivity [W /m°C] 0.041 Speciﬁc heat capacity [J /kg°C] 1450 Table 1: Properties of SPEB’s numerical model. The initial temperature of the model is assumed to be equal to 8°C [22]. Also, the initial stress conditions are obtained from the results of another steady-state simula- tion considering only the gravity e ects. In other words, the ﬁnite element analysis has been done in two stages. Through the ﬁrst stage, a steady state analysis has been carried out while considering the gravity e ects to gen- 7 Figure 3: The thermo-mechanical boundary conditions applied to the model. erate the initial stresses as the input for the transient sim- ulation. In the transient stage, the energy pile response to the applied axial load and thermal cycles has been studied over a ﬁve-year span of time. 3.3. Model Validation To validate the implemented TM model in this work, experimental data studied by Bourne-Webb et al. [14] has been used. The experimental data obtained from the full-scale study that was undertaken on a test pile that was equipped with the heat exchanger pipes. The men- tioned pile was designed as part of a 143 piles founda- tion system of a ﬁve-storey building that was located in the Clapham Centre of Lambeth College London and functioned as a geothermal heat exhchanger of a GSHP system. The pile is installed in the London clay whose properties are reported in Table 2. The conducted tests included both heating and cool- ing cycles in conjunction with a mechanical load. The test pile response to the applied thermo-mechanical loads was obtained using Vibrating Wire Strain Gages (VWSG) and Optical Fibre Sensors (OFS). As shownProperty ConcreteLondon Clay Young’s modulus [Pa] 4010970106 Poisson ratio [-] 0.2 0.236 Unit weight [kN /m3] 20 20 Thermal conductivity [W/m°C]2.33 1.79 Speciﬁc heat capacity [J/kg°C]960 910 Thermal expansion coef- ﬁcient [1 /K]8:51064105 Hydraulic conductivity [m/s]- 11011 Porosity [-] - 0.4 Table 2: Characteristics of the validation model’s material in Fig. 4, the ﬁrst stage of the experiment included ap- plying a 1200 kN mechanical load. During the cooling stage, a thermal load of T=-19C in addition to the 1200 kN mechanical load was applied to the pile. Fi- nally, the heating phase which consisted of the mechan- ical load and a thermal variation of T= +10C was carried out. The initial temperature of the pile and soil in the Lon- don case had been assumed constant over the entire do- main and equal to 19.5°C and the thermo-mechanical 8 loading time applied in this study lasted for almost 1250 days. Figure 4: Thermo-mechanical loading history of experimental case (Bourne-Webb et al. [14]) Comparing the pile head displacements obtained from the experimental tests and the numerical model, good agreement was found during the cooling phase (100 – 860 h) (Fig. 5). Although the magnitude of vertical displacement of the pile head during the heat- ing loading is slightly lower in the numerical model, the similarities are acceptable from an engineering point of view to consider the model an appropriate tool for pre- dicting the thermo-mechanical behavior of the energy pile. Figure 5: Comparison between the numerical and ﬁeld data regarding the vertical displacement of the pile head.Fig. 6 and Fig. 7 provide a comparison between the experimental data which were obtained from two mea- surement systems (VWSG and OFS) and the numeri- cal results in terms of mobilized shaft friction along the pile-soil interface and axial load along the pile axis. A two-tailed student’s t-test is carried out with 95% level of conﬁdence in which a null hypothesis ”Any di er- ences, discrepancies, or suspiciously outlying results are purely due to random and not systematic errors” is made. The p-value which is an outcome of the t-test shows the acceptance or rejection of the null hypothe- sis. For correlation between the numerical and the ex- perimental results with 95% level of conﬁdence, p-value 0.05 is required for every set of data. The obtained p- value for each of the validation graphs are provided in Table 3. The numerical results that were obtained from the TM model showed a good agreement with the ex- perimental data reported by Bourne-Webb et al. [14]. 3.4. Thermo-Mechanical Loading History of the En- ergy Pile in SPEB The validated numerical model in COMSOL Multi- physics is applied to investigate the behavior of one of the geothermal piles of the SPEB’s foundation. The total loading time was 1825 days (ﬁve years) which included the seasonally cyclic thermal loads (i.e., 6 months heating and 6 month cooling per year) in ad- dition to a constant mechanical load. The mechanical load was equal to the ultimate capacity of the friction pile (i.e. 1000 kN) that is applied with a factor of safety of 3. It should be noted that the bearing capacity of the 9 p-valuemobilized shaft friction (VWSG)mobilized shaft friction (OFS)Axial load (VWSG)Axial load (OFS)Pile head dis- placement Mechanical loading 0.134 0.509 0.134 0.101 0.061 Heating loading 0.845 0.447 0.572 0.160 cooling loading 0.837 0.441 0.100 0.577 Table 3: Summary of the p-values obtained from the student’s t-test. (a) Mechanical load (b) Mechanical +Cooling load (c) Mechanical +Heating load Figure 6: Comparison between numerical (Num) and experimental (Exp) results of mobilized shaft friction along the pile-soil interface. (a) Mechanical load (b) Mechanical +Cooling load (c) Mechanical +Heating load Figure 7: Comparison between numerical (Num) and experimental (Exp) results of axial load along the length of the pile. pile is computed by the ultimate end-bearing, and ulti- mate skin friction capacities.As mentioned earlier, the thermal load is applied to the pile while the heat transfer ﬂuid circulates in the heat 10 exchanger pipes. In this work, such a thermal load is represented by a constant temperature, which is equal to the inlet temperature of the heat exchanger ﬂuid, along the length of the heat exchanger pipe. Based on the nu- merical study carried out by Saaly et al. [22], the in- let temperatures of the pipe have been suggested to be equal to 20°C during May–Oct, and -1°C during the 6 cold months (Nov-Apr) of the year. As mentioned be- fore, the inlet temperatures were chosen to meet the thermal balance of the ground during heat extraction and rejection over the year as well as to avoid freez- ing at the soil-pile interface. Fig. 8 provides the sea- sonal variation of the inlet temperature of the heat ex- changer probes as well as the mechanical load applied to the head of the pile. Figure 8: Time histories of applied thermal and mechanical loads. To demonstrate the initial pile settlement due to only the mechanical load, during the ﬁrst 30 days of the sim- ulation the temperature in the probes are equal to the initial temperature ( Tinitial=8C ). In addition, duringthe last 30 days of the loading time history, the tem- perature of the probes are equal to 8C to evaluate the reversibility of pile head displacement induced by the thermal loads. 4. Results and Discussion To study the performance of an energy pile subjected to thermo-mechanical loadings, the vertical displace- ment of the pile, the maximum axial load induced by the thermo-mechanical loads and the mobilized shaft fric- tion along the pile-soil interface are investigated in this section. The reason is that such factors are considered in the Limit State-based geotechnical design. The energy pile and its surrounding soils were sub- jected to an additional thermal load due to the heat loss through the slab of the basement ﬂoor. Such heat dis- sipation from the basement into the ground occurred in spite of insulating the basement envelope [34–37]. The reason was that the basement slab was in direct contact with the foundation soil whose temperature was di er- ent from the regulated temperature of the basement in- door air. Consequently, the thermal gradient between the foundation soil and basement indoor air contributed to heat loss via the thermal conduction through the ﬂoor slab [22, 38]. The impacts of such heat leakage is of- ten ignored while assessing the energy piles behavior. Therefore, this heat loss was incorporated in the model using a convective heat ﬂux which was assigned to the basement slab surface. The energy pile response to the thermo-mechanical loading is provided for two cases: without considering heat loss through the basement slab 11 (Case 1) and with considering the basement energy loss (Case 2). 4.1. Pile Head Displacements The initial pile head displacement that was induced by the mechanical load at the end of the ﬁrst 30 days of the simulation time history was equal to -0.79 mm (Fig. 9). Once heat extraction from the ground by ap- plying a T=9C began, further pile head displace- ment occurred whose maximum reached to -1.19 mm at t=182 days (i.e., the end of the heating season of the ﬁrst year of the system operation). Such an observation was induced due to the contraction of the pile due to the temperature decrease in the probes with respect to the initial temperature of the pile (Fig. 8). For the case with no heat loss through the basement (Case 2), the mini- mum pile head displacement reached -1.18 mm at t = 182 days. This means that during the ﬁrst few months of the analysis, the heat loss through the slab was not ef- fective enough to inﬂuence the pile head displacement. This could be because of the signiﬁcant thermal mass of the ground in the short time. In contrast, during the heating loading (May–Oct), the geothermal pile expanded upon a +12C tempera- ture increase. The maximum expansion, whose value was+0.43 mm relative to the initial settlement, hap- pened at t =1795 days (i.e., the end of the cooling sea- son of the ﬁfth year of analysis). However, without con- sidering the energy loss from the slab, the maximum expansion reached to +0.31 mm which showed occur- rence of 15% less expansion compared to the maximumexpansion of the pile head in Case 2. This discrepancy was due to heat leakage through the basement which increased the soil temperature near the basement ﬂoor and correspondingly caused more expansion at the pile head. It should be noted that the computed pile dis- placements (Fig. 9) were considerably less than the ac- ceptable settlement (i.e., 25 mm) recommended by the National Building Code of Canada [39]. During the last 30 days of the operation of the sys- tem (i.e. at the end of the ﬁfth year; 1795–1825 days), the inlet temperature of the pipes was set equal to the initial temperature of the pile ( Tinitial=8C) to evaluate the reversibility of displacement induced by the thermal loads. It was observed that 40% and 25% of the ini- tial pile head displacement left in the pile for the cases with and without considering heat loss, respectively. In other words, the vertical displacements of the pile head were not fully reversible which is suspected to be due to the elastoplastic behavior of the soil. In addition, for the case in which the basement energy loss was taken into account, the soil temperature increased up to 1.5C near the basement ﬂoor. Therefore, a larger residual dis- placement was observed for this case. 4.2. Axial Load and Mobilized Shaft Friction Variations In Case 1, in the stage in which the pile was subjected to only mechanical load, the pile was fully compressed. However, when a cooling load was applied in addition to the mechanical load to extract heat from the ground (T<0C), the concrete pile contracted. Pile contrac- tion induced tensile forces in the pile. Therefore, the 12 Figure 9: Vertical displacement of the pile head with and without con- sideration of the heat leakage from the SPEB basement. axial load along the pile length decreased compared to the only mechanical loading stage. The maximum re- duction in axial load that was applied on the pile due to the cooling load is 28 kPa (i.e., 9% reduction in the axial load during cooling stage loading compared to mechan- ical loading stage) that occurred at 0.9 m below the pile head (Fig.10a). Contrary, at the end of the heating loading stage, compressive forces were generated in the pile due to the pile expansion. Consequently, up to 3% increase in compressive load compared to the initial mechani- cal load applied on the pile (i.e., 333 kN) was observed along the length of the pile. According to the results regarding the axial load along the pile axis for Case 2, at the end of heating loading stage, the pile was up to 9% more compressed compared to the initial mechanical load applied on the pile (i.e., 333 kN) (Fig. 10b). It was observed that the maximum compressive load in Case 2, which was equal to 363 kN, was induced at the pile depth of 0.6 m dur-ing the heating loading stage. While the maximum ax- ial load that was induced during heating loading at the same pile depth in Case 1 was 342 kN which was 6% lower than the maximum axial load in Case 2. The rea- son was the additional pile expansion in Case 2 due to the heat loss through the basement. Such results signify that considering the buildings heat leakage through the underground enclosure in urban areas may be an im- portant factor in assessing the energy pile response to thermo-mechanical loading. The axial load distribution illustrated in Fig. 10 sig- niﬁed that the thermal cycles change the magnitude of the induced axial load along the pile. In other words, the heating load (i.e., T= +12C) increased the induced axial load along the pile up to 9%. Such an increase in the axial load along the pile should be addressed in the geotechnical design by considering a higher factor of safety. For example, in Case 2 that the obtained maxi- mum compressive axial load along the pile axis was 363 kN, 9% increase in the factor of safety (initially consid- ered to be equal to 3.0) is required to ensure safe ap- plication of the energy pile since the pile was initially designed for allowable load of 333 kN. In practice, a factor of safety of at least 3.5 fulﬁls the required safety margin for the pile that is studied in this work. The distribution of the mobilized shaft frictions along the pile-soil interface under three di erent conditions; (i) mechanical loading, (ii) heating and mechanical loadings, and (iii) cooling and mechanical loadings are provided in Fig. 11. During only the mechanical load- ing, the direction of the skin friction was upward and de- 13 (a) Without considering the heat loss through the basement (b) Considering the heat loss through the basement Figure 10: Axial load distribution along the pile axis. creased by depth. However, during the cooling loading stage (i.e., winter) the mobilized shaft friction along the length of the pile, upper than the neutral point, increased (maximum 3% increase in mobilized shaft friction at the pile depth of 0.3 m compared to the mechanical loading stage) while along the length of pile, lower than the neu- tral point, decreased up to 9% due to the contraction of the concrete pile. Conversely, during the heating load- ing stage that expansion occurred in the pile, the mobi- lized shaft friction increased up to 5% along the length of the pile lower than the neutral point while along the upper part of the pile it decreased (maximum 3% de- crease in mobilized shaft friction at the pile depth of 0.2 m compared to the mechanical loading stage). Comparing the mobilized shaft frictions resulted from Case 1 and Case 2, it is observed that in Case 2 in which heat loss through the basement added more heat-ing load to the ground, shaft frictions decreased up to 4% compared to Case 1. This happened since the addi- tional heating load especially close to the basement ﬂoor induced more expansion compared to Case 1. This ex- pansion led to a downward pile-soil friction which con- tributed to a lower shaft friction along the pile-soil in- terface in Case 2 compared to Case 1. 4.3. Thermal ﬁeld The temperature proﬁles produced by the variation of the inlet temperature of the heat exchanger pipe in Case 1 are illustrated in Fig.12. The temperature ﬁelds of the model are provided for three stages of loading. In the ﬁrst stage, the mechanical loading was only applied to the pile (Fig.12a). Since no thermal load was imple- mented in this stage, the temperature of the all domain was equal to the initial temperature ( Tinitial=8C). At the end of the cooling loading stage 12c, when the 14 (a) Without considering the heat loss through the basement (b) Considering heat loss through the basement Figure 11: Mobilized shaft friction distribution along the soil-pile interface. inlet temperature of the heat exchanger ﬂuid was equal to -1C, the temperature of the pile and the surround- ing soil decreased. Despite the noticeable decrease, the average temperature of the soil at the pile-soil interface was equal to 3C, which is above the freezing point. In other words, the ”non freezing pile-soil interface” crite- rion was met using the inlet temperature of -1C during the cooling loading stage. The temperature ﬁeld at the end of the heating loading stage was shown in Fig. 12c. According to this ﬁgure, the temperature of the soil do- main ranged from 8C to 15C. The temperature ﬁelds of the model for Case 2 are shown in Fig. 13. The top boundary of the axisymmet- ric model in Case 2 was in direct contact with indoor air of the basement whose temperature assumed to be con- stant and equal to 20C. Also, for the side and bottom boundaries of the model, adiabatic condition was con-sidered. At the end of the mechanical loading stage (t =30 days), the soil did not have enough time to absorb heat from the basement ﬂoor. Therefore, the tempera- ture proﬁle of the soil was equal to the initial temper- ature (Fig. 13a). The temperature proﬁle at the end of cooling loading (Nov–Apr), when the inlet temper- ature of the heat exchanger ﬂuid was equal to -1C, is provided in Fig. 13b. According to this ﬁgure, the tem- perature of the soil surrounding the pile decreased up to 5C although above the freezing point. In addition, the soil temperature near the basement slab increased up to 1.5C due to the heat loss through the basement. At the end of the heating loading stage (May–Oct), when the temperature of the heat exchanger ﬂuid was constant and equal to 20C, the temperature of the pile and its surrounding soil increased compared to the ini- tial temperature (Fig. 13c). The temperature of the 15 (a) Mechanical loading (b) Cooling +mechanical loading (c) Heating +mechanical loading Figure 12: Temperature proﬁles at the end of mechanical loading; cooling and mechanical loading; and heating and mechanical loading stages for Case 1. ground far from the e ective diameter of the pile, in- creased up to 1.5C especially at shallower depth of soil. This was expected to happen due to the heat dissipation through the basement of the building. 5. Conclusion In cold regions where the heating load of the build- ing is commonly higher than the annual cooling load, the thermal imbalance may occur in the soil due to an excessive heat extraction from the ground. In design of such systems in cold regions, the energy performance of the energy pile system is of paramount importance. The heat extraction and injection from /to the ground must be designed in such a way to avoid freezing at the pile- soil interface as well as occurring ground thermal im- balance. In addition to the energy performance of the energy pile systems, their TM behavior should be inves- tigated as well. Due to the lack of information regarding the e ect of heat dissipation through the basement slab on the en- ergy pile’s response, a numerical study was carried out.A single geothermal friction pile of an energy foun- dation system that was designed for supplying the en- ergy demand of the Stanley Pauley Engineering Build- ing located in the Campus of the University of Mani- toba, Winnipeg, Canada was considered. To this end, a 2D TM model was implemented in COMSOL Multi- physics and was validated against the ﬁeld data obtained from the in-situ tests in the Lambeth College in London [14]. Based on the numerical analysis, the following conclusions can be made: During the cooling stage , the pile head downward displacement increased due to pile contraction. Al- though noticeable, the pile displacements were be- low the allowable pile settlement according to the National Building Code of Canada [39]. The results veriﬁed that when the energy pile was subjected to heating and cooling loads, the thermal deformations generated additional com- pressive and tensile stresses during heating stage (May–Oct) and cooling stage (Nov–Apr), respec- tively. This necessitates further considerations in 16 (a) Mechanical loading (b) Cooling +mechanical loading (c) Heating +mechanical loading Figure 13: Temperature proﬁles at the end of mechanical loading; cooling and mechanical loading; and heating and mechanical loading stages for Case 2. the geothechnical design and application of energy pile systems in cold regions. These considerations can be implemented in the form of increasing the factor of safety e ective on the allowable load. The obtained results showed the importance of considering heat dissipation through the basement in evaluating the thermo-mechanical behavior of energy piles. The reason was that the heat loss through the basement worked as an additional ther- mal load which intensiﬁed the heating load and weakened the e ects of applied cooling load on the pile. As the thermal loads were applied to the pile, the mobilized shaft friction along the pile-soil inter- face varied depending on the type of the loading stage (i.e., heating or cooling). For example, heat- ing load caused up to a 7% decrease in shaft fric- tion along the pile upper than the neutral point compared to shaft frictions induced during the me- chanical loading stage. The same thermal load de- creased the shaft friction along the pile lower thanthe neutral point up to 12%. This work provided an insight into the mechanical be- havior of a single pile of a energy foundation system. Result of this study in conjunction with thermal perfor- mance analysis can facilitate further application of such systems in cold regions and more speciﬁcally Canada. 6. Acknowledgment The authors gratefully acknowledge the ﬁnancial sup- port of the NSERC CREATE SERA Program and the University of Manitoba. Also, the authors would like to acknowledge the Faculty of Engineering of the Uni- versity of Manitoba for the technical support of this re- search.