Experimental and numerical study on a novel low temperature façade solar thermal collector to decrease the heating demands: A south-north pipe-embedded closed-water-loop system

Solar Energy 147 (2017) 22–36

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Experimental and numerical study on a novel low temperature façade solar thermal collector to decrease the heating demands: A south-north pipe-embedded closed-water-loop system Mohamad Ibrahim a,b,⇑, Etienne Wurtz a,b, Jocelyn Anger a,b, Oussama Ibrahim c a b c

Univ. Grenoble Alpes, INES, F-73375 Le Bourget du Lac, France CEA, LITEN, Department of Solar Technologies, F-73375 Le Bourget du Lac, France Faculty of Engineering, Lebanese University, Beirut, Lebanon

a r t i c l e

i n f o

Article history: Received 20 October 2016 Received in revised form 14 February 2017 Accepted 21 February 2017

Keywords: Active pipe-embedded wall Façade solar thermal collector Energy efficiency Low-grade energy source Thermally-active envelope

a b s t r a c t Recently, more and more research is being conducted on thermo-activated building walls with the use of circulating water or other fluids for the aim of decreasing and shifting the heating and/or cooling loads. In this context, we present a novel concept of the active embedded-pipe envelope systems. The system consists of an active closed-loop-water-pipes embedded in the building exterior walls to harvest and utilize the solar energy gain on the south wall exterior surface to decrease or offset the heat loss through the north wall and enhance thermal comfort. During non-cloudy winter days, a significant amount of solar energy hitting the south (insulated) facade is not transferred to the inside environment. In this study, a comparative experimental set-up is carried out to test the system’s efficiency and compare its performance to that of a static insulated envelope without the system. Also, a numerical model is developed and validated against experimental measurements. Numerical simulations with a parametric study are carried out to examine the active wall loop system’s efficiency for different design and operating parameters. The main conclusion derived from this study is that the system performs very well in the Mediterranean climates (or similar climates) and to a lower extent in the cooler ones. However, its performance is highly dependent on several design, climatic, and operating variables which should be optimized to have the best performance. Ó 2017 Elsevier Ltd. All rights reserved.

1. Introduction The building fabric plays a major role in regulating the indoor environment. Through influencing the energy flows between the indoors and the outdoors, the building walls are key components for energy saving in the building enclosure structure. Traditionally, the building envelope has been treated as a static or passive system in the building thermal energy system allowing or resisting heat energy transfer from the inside to the outside or vice versa. However, as our needs have evolved and technologies have advanced, the demand placed on designers to integrate a wide range of increasingly complex materials, components, and systems into the building enclosure has grown. One of the methods is to use low-grade energy sources such as solar energy, underground water, cool air, and geothermal energy ⇑ Corresponding author at: CEA, LITEN, Department of Solar Technologies, INES, F-73375 Le Bourget du Lac, France. E-mail address: [email protected] (M. Ibrahim). http://dx.doi.org/10.1016/j.solener.2017.02.036 0038-092X/Ó 2017 Elsevier Ltd. All rights reserved.

in an active embedded-pipe envelope system that utilizes the circulating water to transfer heat to or from the inside space. This is quite popular for the ceiling and floor, such as chilled ceiling systems and under-floor heating systems. A comprehensive review of the research and application of active hollow core slabs in building systems for utilizing low energy sources is presented in Xu et al. (2014) and Xu et al. (2010). Fewer studies have dealt with embedded-pipe systems in the building exterior fabric, particularly the exterior walls, with water or fluid circulating inside for heating and cooling. As reported by Yu et al. (2016), this approach can work with lower temperature gradients, leading to relatively easier utilization of low-grade energy sources and consuming less energy compared to air-based systems. Recently, more and more research is being conducted on thermo-activated building walls with the use of circulating water or other fluids for the aim of decreasing and shifting the heating and/or cooling loads. In a very recent study, Yu et al. (2016)

M. Ibrahim et al. / Solar Energy 147 (2017) 22–36

proposed a mini-tube embedded capillary-network and thermalwater activated building envelope to turn the exterior wall passive component into an active one. This system can act as a thermal barrier system and might also supplement cooling or heating energy to the space depending on the temperature of circulated water. Based on the results generated from a dynamic simulation model for summer conditions, the authors showed that the thermo-activated wall can be effective in stabilizing the internal surface temperature, offsetting the heat gain, and supplying cooling energy to the space. This system was further investigated by examining the effect of three locations inside the wall including external, middle and internal side on inner surface temperatures and energy consumption (Niu and Yu, 2016). The results indicated that the internal wall surface temperature can be neutralized from the ambient environment when the embedded tubes are fed with thermal water. The wall can work with a wide range of water temperature and the optimal location of the tube network is relatively constant in different modes. Xie et al. (2015) conducted an experimental study on a pipeembedded building exterior wall to validate a numerical frequency-domain finite-difference model (Zhu et al., 2014). An experiment test rig was developed for the measurements of the thermal responses of the pipe-embedded building envelope under pre-defined dynamic conditions. During summer weather conditions, good heat intercept effect and space conditioning potential of the pipe-embedded building envelope were observed from the measurements, according to the authors. When the supply water temperature is 17.5 °C, the heat absorption of the internal surface of the pipe-embedded building envelope may reach up to 25 W/ m2. This value will be further increased by changing the position of the embedded pipe more close to the internal surface of the pipe-embedded building envelope. Another system is proposed by Zhang et al. (2014) to decrease the heating and cooling demands. They studied the use of heat pipes implanted in the exterior walls. It can transfer a large amount of heat without additional power consumption when the temperature difference is small (Noie, 2005). The heat transfer performance and energy-saving characteristic were investigated theoretically and experimentally during winter conditions with the typical meteorological data of Jinan, China. The results showed that the heat loss of the south external wall is reduced by 14.47% during the winter in a typical year. In this study, we present a novel concept of the active embedded-pipe envelope systems. The system consists of an active closed loop water pipes embedded in the building exterior walls to harvest and utilize the solar energy gain on the south wall exterior surface to decrease or offset the heat loss through the north wall. During non-cloudy winter days, a significant amount of solar energy hitting the south, usually insulated, facade is not transferred to the inside environment. Due to the high thermal resistance of the insulation, the majority of this energy is dissipated to the outside. In the following sections, we present (1) the system’s concept, (2) a comparative experimental study to test the performance of the proposed system, (3) a numerical model with the simulation results compared to the experimental measurements, and (4) the simulation results for some case studies for different climates, along with a parametric study.

2. The south-north wall closed water-loop system The active embedded-pipe south-north closed-water-loop system is shown in Fig. 1. The exterior walls are composed of concrete or brick layer with an outside aerogel-based insulating rendering.

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The latter has a thermal conductivity of 0.026–0.027 W/(m K) (Ibrahim et al., 2015a,b). It is applied on the building facades through spraying using a plastering machine. Its application through spraying facilitates the implementation of such envelope embedded-pipe systems. In the south wall, the water pipes are placed at just a few millimeters from the exterior surface of the insulating rendering as a serpentine shape. In the north facade, they are placed in direct contact with the concrete just at the interface between the insulating rendering and the concrete layer. By means of a pump, water circulates through the south facade to collect the heat energy provided by solar radiation. It then passes through the east or west facade to reach the north one where it dissipates its heat through the concrete layer to the inside. Also, some of the energy is stored in the thermal mass (concrete) to be used in a later time of day. Finally, the water exiting from the north facade returns to the south one. For the east or west walls, the water pipe can be placed at the middle of the insulating rendering or in direct contact with the concrete layer.

3. Experimental study 3.1. Test units description Two experimental units having the same dimensions and construction materials have been established at the INES (national institute of solar energy) experimental field near Chambery, France (latitude 45.65°N, longitude 5.86°E). The first one serves as the test experimental unit with the active embedded pipes south-north wall loop system being implemented and the other serves as the reference one with no fluid pipes (see Fig. 2). Each experimental unit has a length of 2.25 m, a width of 1.6 m and a height of 1.2 m. The walls are composed of 12 cm of concrete layer with 4 cm of aerogel-based rendering. In the south facade, the fluid pipes are embedded in the aerogel rendering and are placed near the exterior surface. These pipes are hold through metallic bars with horizontal fixings as shown in Fig. 3 (upper part) that are fixed to the concrete layer through bolts. In the north facade, the pipes are in direct contact with the concrete layer. Since the aerogel rendering is highly insulating, and to enhance the heat transfer from the fluid pipes in the north facade to the interior space, a layer of traditional plastering is added between the concrete and the aerogel rendering which covers the pipes (see Fig. 3 lower part). The ground and ceiling are composed of a layer of thermal insulation to limit the heat transfer from these surfaces. A final finishing having a solar absorptivity of 0.6 is applied on the exterior surfaces of all the facades. The pipe’s diameter is 1.2 cm and the pipe’s spacing is approximately 10 cm. The system is equipped with a variable flow rate pump connected to the pipes with an expansion vessel for pressure regulation. The pump’s flow rate could be varied from 5.53 L/h to 116 L/h. Also, the pump is equipped with a control timer that sets the pump ‘‘ON” or ‘‘OFF” during pre-assigned scheduled hours of the day. The heat transfer fluid is a mixture of water/anti-freeze (60%/40%) to protect the system when exposed to low temperatures. Temperature sensors (Type T thermo-couples) are placed at different positions to measure the wall and fluid temperatures. For each of the south and north walls, three sensors are placed at the exterior surface and other three sensors are placed at the interior surface at different heights: near the bottom (at 32 cm from the bottom), at the middle, and near the top edge (at 92 cm from the bottom). For each of the other walls and the roof, one sensor is placed at the middle of the exterior surface and another one at the middle of the interior surface. Also, a temperature sensor is placed at the inner surface of the floor. Other than the wall surface

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Fig. 1. The south-north active embedded-pipe wall closed loop system (left: south façade, right: north façade).

Reference unit Test unit

(a)


Direct normal solar radiation via a pyrheliometer mounted on a solar tracker (Kipp and Zonen CH1 pyrheliometer Kipp, 2014b). The latter works in the range of 0–4000 W/m2 with an accuracy of ±10%.
Wind direction and speed via an anemometer station (Windsonic WS1 anemometer Windsonic, 2014) that has a measuring range of 0–60 m/s with a precision of ±2 m/s at 12 m/s for the wind speed and a measuring rage of 0–359 °C with a precision of ±3° at 12 m/s for the wind direction.
Ambient air temperature and relative humidity using a hygrothermal sensor (Campbell scientific CS215 Campbell Scientific, 2014) placed in a multi-plate solar shield. The latter has a temperature measurement range of 40 °C to +70 °C with a precision of ±0.4 °C for temperature ranges from +5 °C to +40 °C and ±0.9 °C otherwise. For the relative humidity, it has a measurement range of 0–100% with an accuracy of ±2% for the range between 10 and 90%; otherwise, the accuracy is ±4%. 3.3. Experimental results

(b) Fig. 2. The experimental test and reference units (a) before and (b) after applying the aerogel-based rendering on the facades and the insulation on the top.

temperatures, the fluid temperature is measured. For each of the north and south facades of the test unit, the inlet and the outlet temperatures are measured as well as the temperature at a position near the middle of the serpentine loop. In addition, a PT100 temperature sensor is placed at the middle of each of the units to measure the interior air temperature. 3.2. On-site weather station A weather station is installed at the experimental field to measure all the environmental inputs. These include the solar radiation, air temperature and relative humidity, and wind speed and direction. The measured data are recorded every minute. The measured climatic variables are the following:
Global solar radiation on a horizontal plane and the global solar radiation on the south facade measured via two pyranometers (Kipp and Zonen CMP22 and CMP11 Pyranometers Kipp, 2014a). The pyranometers work in the range of 0–4000 W/m2 with an accuracy of ±10%.

3.3.1. Results of the facade solar collector test unit In this section, we are focusing on the thermal behavior and performance of the test unit in terms of the amount of collected heat flux and the daily efficiency of the facade solar collector during different days of the testing period. In the following section, the comparison of the surface and air temperatures is carried out between the test unit and the reference unit. Measurements are carried out for several months during the winter season, from November 2015 till March 2016. Measurements are recorded every 2 min. The air temperature in the test units is free-running (no heating system). The pump is set ‘‘ON” from 9 a.m. till 5 p.m. Fig. 4 shows the outlet fluid temperature of the south wall, inlet fluid temperature of the south wall, the outside dry-bulb temperature (To), and the pumps control (ON/OFF) and the south solar irradiation (I_G) and the collected heat flux (Qcollected) at the south wall, during 24 days in November. The collected heat flux, Qcollected, is calculated using Eq. (1).

_ fluid  cpfluid  DT fluid Q collected ¼ m

ð1Þ

_ fluid ; cpfluid ; and DT fluid are the fluid’s mass flow rate, specific where m heat, and the fluid’s temperature difference between the outlet and the inlet. As highlighted by several authors (Srivastava et al., 1982; Nayak et al., 1989; Bopshetty et al., 1992; D’Antoni and Saro, 2012), the notion of the instantaneous energy efficiency becomes meaningless in the case of massive solar thermal collectors, because inertial effects delay the energy response at time instants where no availability of solar radiation is present. This leads to the conclusion

M. Ibrahim et al. / Solar Energy 147 (2017) 22–36

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Fig. 3. The pipes position and fixing in the south (left) and the north (right) facades.

Fig. 4. The test unit’s outlet fluid temperature of the south wall, inlet fluid temperature of the south wall, the outside dry-bulb temperature (To), the pumps control (ON/OFF), the south solar irradiation (I_G), and the collected heat flux during November 2015.

that under certain circumstances (e.g. cloud passing) the energy efficiency would be very close to 100%. Thus, some authors refer to medium (daily or weekly) or long-term (season) energy efficiency. The latter is the ratio of the collected heat energy during a certain period of time to the available solar energy during the  day , is used and is same period. In this study, the daily efficiency, g calculated using Eq. (2).

g day ¼

R day 0

A

Q collected R day Isolar 0

ð2Þ

where A; and Isolar are the collector (facade) area and south solar irradiation, respectively. During the days of high solar radiation, for e.g. the 5th day till the 9th day where the south solar irradiation reaches a max-

imum value of around 900 W/m2, the outlet fluid temperature from the facade solar collector (south wall) reaches values of more than 30 °C during the day, and the fluid is heated up to 10–11 °C. The collected heat flux during these days reaches a maximum value in the range of 200–250 W/m2. The daily efficiency of the proposed system ranges between 18% and 21% with a daily energy harvesting between 0.95 kW h/m2 and 1.26 kW h/m2. During the days with medium solar irradiation (maximum is around 350–400 W/m2 during most of the day times), for e.g. from the 11th till the 14th days, the fluid is heated up to 6–7 °C during day time. With inlet temperatures of around 14–16 °C, the outlet fluid temperatures are in the range of 21–24 °C. The daily collected heat energy is between 0.18 kW h/m2 and 0.29 kW h/m2.

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During the days with low solar irradiation, for e.g. the 20th and the 21st days, where the solar irradiation stays below 150 W/m2 most of the day time, the outlet fluid temperature is very similar to the inlet fluid temperature with no significant possibility for energy harvesting. Fig. 5 shows the same parameters as Fig. 4 but from the 17th of December 2015 till the 10th of January 2016. The difference between these two periods is the outside dry-bulb temperature values. In the previous period (Fig. 4), the daily temperature varies between a maximum value of around 20 °C at day-time and a minimum value of around 5 °C during the days with high solar irradiation. In the second period (Fig. 5), the outside air temperature reaches a maximum value of around 13 °C during the days with high solar irradiation with a minimum value that goes below 0 °C at night. Considering, the days with high solar radiation (for e.g. 25/12, 26/12, 27/12, 28/12), the collected heat energy of the facadesolar collector is around 0.98 kW h/m2 to 1.17 kW h/m2 per day, corresponding to a daily efficiency of around 22%. With inlet fluid temperatures to the south wall of around 15 °C, the outlet fluid reaches a temperature of around 25 °C. 3.3.2. Results comparison between the test and the reference units Fig. 6 shows the internal surface temperature at three heights (bottom, middle, and top) of the north facade for the test unit and the internal surface temperature (middle position/height) of the reference unit along with the exterior air temperature and the south solar radiation. It is shown that the north inner surface temperatures of the test case (with the embedded-pipe loop system) are significantly higher than those of the reference case. By comparing the temperatures at the middle of the interior surface, the difference between the two can reach more than 4.5 °C for the days with high solar irradiation (for e.g. 05/11 till 10/11). For the days 3/11 and 4/11, the surface temperature decreases compared to the previous day due to the relatively lower solar radiation but it remains higher than that of the reference case. During a specific day, and due to the effect of thermal storage in the concrete structure, the north wall inner surface temperature reaches its maximum value at approximately 7 p.m. then starts to decrease gradually where the wall dissipates its energy to the inside space. Due to this storage, we can see that the temperature decreases slowly at night providing heat to the space. Also, a significant time shift, of 6–7 h, is observed between the inner surface temperature and the exterior surface temperature. The temperature of the air inside the test and the reference units are shown in Fig. 7. It is illustrated that the air temperature of the test case is always higher than that of the reference case.

The difference is much higher during the days with high solar irradiation where it reaches almost 3 °C. Fig. 8 shows the south wall exterior and interior surface temperatures for both the test and reference units. It is shown that during day time, the exterior surface temperatures of the reference unit are higher than those of the test case especially during the days with high solar irradiation. The difference between the two during these days can reach up to 6–7 °C. The reason that explains this is the presence of the fluid pipes near the outside surface of the south facade (in the test unit) which receives part of the solar energy and transfers it to the north façade. This cools down the outside surface temperature of south wall. For the interior surface temperature, the temperatures of the two units are very close, except in the days with high solar irradiation (for e.g. from 6/11 till 10/11) where the inner surface temperatures of the test unit are higher than those of the reference unit, although it is the opposite case for the exterior surface. The reason that can explain this is the heat gain received due to the long-wave heat exchange with the relatively ‘‘hotter” north wall (of the test unit compared to the reference one) and the convective heat exchange with the relatively ‘‘hotter” inside air. The previous period in November has high solar irradiation with relatively high outside air temperature during day time. To examine the systems performance during a longer period, the interior surface temperature of the north wall of the test unit (at middle position) and of the reference unit as well as the south solar irradiation and the outside dry bulb temperature from the 17th of December 2015 till the end of February 2016 are shown in Fig. 9. The same conclusions can be drawn as those presented for the period of November. The north wall inner surface temperatures of the test unit are significantly higher than those of the reference unit (more than 4 °C difference for the days with high solar irradiation). 4. Numerical model 4.1. Mathematical equations The heat transfer within the wall structure embedding the pipes is a complex 3-dimensional problem. In our case here, we simplified the heat transfer into 2-dimensional which is coupled to the moving fluid in the water pipes. Some assumptions are made:
The serpentine water loop is modeled as a long horizontal pipe.
Symmetry boundary conditions are considered between the serpentine pipes.

Fig. 5. The test unit’s outlet fluid temperature of the south wall, inlet fluid temperature of the south wall, the outside dry-bulb temperature (To), the pumps control (ON/OFF), the south solar irradiation (I_G), and the collected heat flux during December and January.

M. Ibrahim et al. / Solar Energy 147 (2017) 22–36

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Fig. 6. The internal surface temperature at three heights (bottom, middle, and top) of the north facade for the test unit and the internal surface temperature (middle) of the reference unit along with the exterior air temperature and the south solar irradiation for 3 weeks in November.

Fig. 7. The inside air temperature of the test and the reference units.

Fig. 8. The south wall exterior and interior surface temperatures for both the test and the reference units.


The circular shape of the pipe is modeled as a square with the pffiffiffiffi same area. Thus, the side of the square has a length of p r, where r is the radius of the pipe.
Heat conduction in the longitudinal pipe direction is neglected since the conductive heat transfer is small compared to the fluid enthalpy flow.

The pipe is divided into several segments. Fig. 10 shows the face view of the south facade and the side view of one segment. For each segment, the 2D heat transfer equation is solved. The domain is discretized into rectangular control volumes as shown in Fig. 10. The outlet fluid temperature from the previous segment is an inlet temperature of the fluid in the current segment.

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M. Ibrahim et al. / Solar Energy 147 (2017) 22–36

Fig. 9. The internal surface temperature (middle) of the north facade for the test unit and the internal surface temperature (middle) of the reference unit along with the exterior air temperature and the south solar irradiation for the period from 17th of December 2015 till the end of February 2016.

Fig. 10. South wall face view (left) and section view (right).

For the wall’s control volumes, the energy balance is written as:

qi;j cpi;j

X @T i;j U adj ðT adj  T i;j Þ dxi;j dzi;j ¼ @t adj

ð3Þ

where the subscript (i, j) represents a control volume. The parameters q, cp, dx, dz, T, and t are the density, specific heat, control volume width, control volume height, temperature, and time, respectively. U is the heat transfer coefficient between two adjacent control volumes and is calculated as following (taking the heat transfer coefficient with the west control volume, Uw, as an illustration):

8 > U ¼ > > < w > > > : Uw ¼

dzi;j dxi;j dx þ i1;j 2ki;j 2ki1;j

;

dzi;j corr dxi;j dx þ 1 þ tube 2ki;j hw;c ktube

for nodes not adjacent to the water node ; for nodes adjacent to the water node

ðaÞ ðbÞ ð4Þ

where k is the thermal conductivity, corr is a correction factor to account for the difference in the lateral area between the modeled square and the real circular shape of the pipe; it is given in Eq. (5). This modeling approach of the pipe with the correction factor is adopted in the TRNSYS Type 360 (FORT, 2001).

pffiffiffiffi Lateral Areacircle 2pr p ffiffiffi ffi p corr ¼ ¼ ¼ 2 Lateral Areasquare 4 pr

ð5Þ

The energy balance for the fluid control volume is shown in Eq. (6):

qf cpf

X @T f dT _ f f U adj ðT adj  T f Þ  mcp dxf dzf ¼ @t dy adj

ð6Þ

_ and hw,c are the fluid mass flow rate and the convective where m heat transfer coefficient between the fluid and the pipe. The last term in Eq. (4) corresponds to the enthalpy change due to fluid flow. The heat transfer at the outside and inside wall surfaces are given in Eqs. (7-a) and (7-b), respectively.

hout ðT out  T x¼0 Þ þ a  I þ hrsky ðT sky  T x¼0 Þ þ hrgr ðT gr  T x¼0 Þ ¼ k

@Tð0; tÞ @x

hin ðT x¼L  T in Þ þ

ð7-aÞ S X @TðL; tÞ hr s ðT x¼L  T s Þ ¼ k @x s¼1

ð7-bÞ

where hout and hin are the outside and inside convective heat transfer coefficients, respectively. hrsky and hrgr are the radiative heat transfer coefficient between the outer wall surface and the sky and the ground, respectively. hrs is the radiative heat transfer coefficient. a and I are the surface solar absorptivity and solar irradiation, respectively. T out ; T in ; T x¼0 ; T x¼L ; and T s are the outside air temperature, the inside air temperature, the wall’s outside surface

M. Ibrahim et al. / Solar Energy 147 (2017) 22–36

temperature, the wall’s inner surface temperature, and the inner surface temperature of the other surfaces enclosing the room. The heat transfer coefficient between the fluid and the pipe wall when the fluid is in flowing state is given in Eq. (8).

hw;c ¼

Nu  kf D

ð8Þ

where kf, D, and Nu are the fluid thermal conductivity, pipe diameter, and the Nusselt number, respectively. The latter is calculated using Eq. (9) (Cengel, 2005)

(

Nu ¼ 3:66; Nu ¼ 0:023  Re

for laminar flow 0:8

1 3

 Pr ; for turbulent flow

ð9Þ

where Re and Pr are the Reynold and Prandtl numbers, respectively. These are calculated using Eqs. (10) and (11).

Re ¼ Pr ¼

V D

v

v

ad

ð10Þ ð11Þ

where V, D, v, and ad are the fluid velocity, pipe’s diameter, kinematic viscosity, and thermal diffusivity, respectively. Concerning the fluid in stagnation mode, it is hard to find correlations for the heat transfer coefficient. Martini and Churchill (1960) studied the natural convection for air inside a horizontal cylinder and concluded that the convective heat transfer for air inside a 10 cm cylinder corresponds to a coefficient of approximately 1.7 W/(m2 K) over a wide range of wall-temperature difference (small as well as large temperature difference). Li et al. (2015) modeled the convective heat transfer coefficient inside horizontal tubes when the water is in stagnation mode as purely conductive. In this study, we tried different values for the heat transfer coefficient between the fluid and the pipe wall when the fluid is in stagnation state, including the purely conductive approach, and it was found that a value of 2 W/(m2 K) gives satisfactory results when compared with the experimental data as shown here after when calculating the statistical indicators in Table 1. It is to be noted also that initial simulations are carried out for different values of the heat transfer coefficient when the fluid is in stagnation state and the results (which are not shown for brevity in the simulation section afterwards) showed that this parameter has no noticeable effect on the system’s performance. 4.2. Model validation The results of the numerical model are compared to the experimental measurements of the test unit with the embedded pipes. The measured south solar radiation, the outside air temperature, the inside surface temperatures of the east wall, west wall, ground and ceiling are inputs to the numerical model. The simulation time-step is taken as 2 min corresponding to the measured data. After doing a mesh refinement test, the calculation domain is divided into a total of 600 control volumes. For the north wall, the outlet fluid temperature and the fluid temperature at the middle of the serpentine pipe resulting from the model simulation are compared to those of the experimental data, and are shown in Fig. 11. Also, the inner surface temperatures at the three different heights (bottom, middle, and top) are compared to those of the experiment. To have a ‘‘qualitative” assessment of the agreement between the simulation and the measurements, the simulated and the measured temperatures are shown in Fig. 12 for the north wall inner surface temperature at the (a) top position, (b) middle position, and (c) bottom position.

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It is illustrated from Figs. 11 and 12 that the model is capable of correctly predicting the water and the wall surface temperatures. However, there is a little over-estimation of the highest values of the water temperature predicted by the numerical model. Also, the water temperatures of the model reaches lower temperatures at night than those measured at the middle position of the serpentine pipe, but they stay higher (than those measured) at the outlet of the serpentine pipe. Concerning the wall inner surface temperatures, the analysis of the time profiles reveals that the accuracy is slightly worst during the decline period (decreasing temperatures curves) where the measured wall temperatures decrease in a faster way than the simulated ones. In addition, the minimum of the simulation curves remains higher than that of the measured curve. As shown in Fig. 12(c), the model has a tendency to overestimate the surface temperature at the bottom position for almost all the considered period. On the contrary, during the increasing temperature phase, a better agreement between the experimental data and the model results is achieved for both the water temperatures and the wall surface temperatures at all positions. To give a ‘‘quantitative” estimation of the agreement between the measured data and the simulated values, the following indicators are calculated:
The percentage root mean square error (PRMSE), which is defined as

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  2 1 Xn Tsi  Tmi PRMSE ¼ i¼1 n Tmi

ð12Þ

where n is the total number of measurements over a certain period and Tsi and Tmi are the simulated and the measured temperature values, respectively.
The root mean square error (RMSE), which is calculated using Eq. (13)

RMSE ¼

rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 Xn ðTsi  Tmi Þ2 i¼1 n

ð13Þ


The average value of the absolute difference between the simulated and the measured values (Eq. (14)), together with its standard deviation (Eq. (15)):

j ¼ jD

1 Xn jTsi  Tmi j i¼1 n

rjD j ¼

rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 Xn  jÞ2 ; with jDj ¼ jTsi  Tmi j ðjDji  jD i i¼1 n

ð14Þ

ð15Þ


The maximum difference between the daily maximum simulated temperature and the daily maximum measured temperature as well as the mean of these maximum differences, which using  are calculated   Eqs. (16) and (17) 

Dmax;peaks  ¼

max Tsmax;day  Tmmax;day 

day¼1 to N

N  X     max;peaks  ¼ 1 D Tsmax;day  Tmmax;day  N day¼1

ð16Þ

ð17Þ

where N is the number of days, Tsmax and Tmmax are the simulation and experimental maximum daily temperature.
The maximum difference between the daily minimum simulated temperatures and the daily minimum measured temperatures as well as the mean value of these minimum differences, which are calculated using Eqs. (18) and (19)

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M. Ibrahim et al. / Solar Energy 147 (2017) 22–36

Fig. 11. The experimental data and the numerical model results of (a) the north wall outlet fluid temperature and of (b) the fluid temperature at a middle position of the serpentine pipe.

Fig. 12. The experimental data and the numerical model results of the north wall inner surface temperature at the (a) top, (b) middle, and (c) bottom height positions.

jDmax;troughs j ¼  max;troughs j ¼ jD

  max Tsmin;day  Tmmin;day 

day¼1 to N

N   1 X Tsmin;day  Tmmin;day  N day¼1

ð18Þ ð19Þ

The PRMSE and RSME testify the good agreement between the simulated values and the measured data. As shown in Table 1, for the north wall, the PRMSE of the water temperature is around 8.9% and 12.3% for the outlet and the middle positions, respectively. As for the inner surface temperature, the PRMSE is around 5.9% for the top and middle positions and around 7.5% for the bottom position. The average difference between the measured and  jÞ is around 0.9 °C for the water temperathe simulated values (jD tures with a standard deviation of around 0.6 °C, and around 0.65 °C for the inner surface temperatures with a standard deviation of around 0.37 °C.

The maximum difference for the daily peak temperatures between the simulated and the measured values doesn’t exceed 1.26 °C, with an average that doesn’t exceed 0.58 °C. However, the difference is more pronounced for the temperature troughs. The maximum difference is around 2 °C for the water temperatures (with an average of around 1 °C) and around 1.2 °C for the surface temperatures (with an average of around 0.7 °C). The same figures are, approximately, applicable to the south wall. The measured and simulated daily collected heat energy (in kW h) over the considered period is shown in Fig. 13. This figure reveals that the numerical model underestimates the collected heat energy for almost all days. The maximum absolute difference between the two doesn’t exceed 0.22 kW h for all days. For the days with, relatively, high energy harvesting, the relative difference ranges between 5% and 10% with an average of 7.3%. For the

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M. Ibrahim et al. / Solar Energy 147 (2017) 22–36 Table 1 The performance indicators to assess the simulation values agreement with the experimental data. PRMSE

RMSE

j jD

rjD j

|Dmax,peaks|

 max;peaks j jD

|Dmax,troughs|

 max;troughs j jD

North wall Tw_out Tw_mid Tint_top Tint_mid Tint_bot

12.29% 8.89% 5.94% 5.96% 7.48%

1.19 °C 0.99 °C 0.63 °C 0.67 °C 0.79 °C

0.99 °C 0.84 °C 0.53 °C 0.57 °C 0.69 °C

0.67 °C 0.55 °C 0.35 °C 0.35 °C 0.40 °C

1.09 °C 1.21 °C 0.81 °C 0.96 °C 1.26 °C

0.37 °C 0.50 °C 0.28 °C 0.35 °C 0.58 °C

2.18 °C 1.99 °C 1.26 °C 1.18 °C 1.23 °C

1.15 °C 1.06 °C 0.81 °C 0.72 °C 0.68 °C

South wall Tw_out Tw_mid Tint_top Tint_mid Tint_bot

13.59% 10.1% 6.68% 6.13% 8.32%

1.56 °C 1.21 °C 0.84 °C 0.70 °C 0.94 °C

1.21 °C 0.97 °C 0.69 °C 0.61 °C 0.74 °C

0.71 °C 0.66 °C 0.44 °C 0.35 °C 0.49 °C

2.1 °C 1.74 °C 1.32 °C 1.12 °C 1.44 °C

0.63 °C 0.57 °C 0.53 °C 0.41 °C 0.61 °C

1.77 °C 1.87 °C 1.12 1.43 °C 1.21 °C

1.21 °C 1.22 °C 0.22 °C 0.66 °C 0.70 °C

Fig. 13. The measured and simulated daily collected heat energy.

days with medium energy harvesting, the relative difference is inherently higher and it is around 15% for all these days. The analysis of the daily energy shows that the numerical tool is able to estimate, to a large extent, the amount of energy harvested through the water circulation in the wall, but there is a tendency in the model to underestimate this energy.

Table 2 Base case design parameters.

5. Case studies and simulation results 5.1. Base case description The numerical model is used to assess the system’s performance over longer periods of time for different climates. The system’s performance is compared to the static insulation system for the same wall structure (insulating rendering outside layer and concrete inside layer) but with no fluid pipes embedded in the envelope. The base case design is shown in Table 2. Simulations are carried out for the base case (with a 10 min time step), then a parametric study is realized to investigate the effect of the some design parameters on the system’s performance and efficiency. The system’s performance is examined for different cities: Barcelona (latitude 41°230 N, longitude 2°110 E), Marseille (43°170 N, 5°220 E), Milan (45°270 N, 9°110 E), and Zurich (47°220 N, 8°320 E). The source of the weather files is the IWEC (International Weather for Energy Calculations) (ASHRAE, 2001). The IWEC data files are typical weather files suitable for use with building energy simulation programs. These are derived from up to 18 years of archived hourly weather data. The simulations are carried out for 6 months, from the beginning of October till the end of March.

Design parameter

Value

South facade thermal collector area North facade sink area Pipe’s diameter Pipe spacing Pipe material Exterior surface solar absorptivity South insulating rendering thickness North insulating rendering thickness Concrete thickness Concrete density Concrete specific heat Concrete thermal conductivity Fluid mass flow rate Fluid mixture Control strategy

4 m2 4 m2 12 cm 12 cm Copper 0.6 7 cm 7 cm 15 cm 2000 kg/m3 1000 J/(kg K) 1.7 W/(m K) 0.011 kg/s Water/anti-freeze (60%/40%) Pump ‘‘ON” if solarsouth > I_controla (W/m2) 19 °C (all time)

Inside Temperature control a

I_control is variable for the different cases.

Different studies used different values for the outside and inside heat transfer coefficients. The outside convective heat transfer coefficient usually ranges between 12 W/(m2 K) and 25 W/(m2 K) as considered in many studies. For example, hout was taken as 15 W/(m2 K) in Yang and Li (2008), 16.2 W/(m2 K) in Daouas (2016), 16.67 W/(m2 K) in Kontoleon and Bikas (2007), 17 W/ (m2 K) in Al-Khawaja (2004), 22 W/(m2 K) in Ozel (2016), and 23 W/(m2 K) in Niu and Yu (2016). In this study, hout is taken as 17 W/(m2 K).

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For the inside heat transfer coefficients, most of the studies consider a value between 6 W/(m2 K) and 9 W/(m2 K) for the combined convective and radiative coefficient. It was taken as 6.2 W/ (m2 K) in Daouas (2016), 6.5 W/(m2 K) in Yang and Li (2008), 8 W/(m2 K) in Niu and Yu (2016), and 9 W/(m2 K) in Ozel (2016). In this study, the inside convective coefficient is taken as 3 W/ (m2 K) according to ASHRAE Handbook (ASHRAE, 1985). The radiative heat transfer coefficient is considered as 4.5 W/(m2 K) which is the value used in the WUFI software (IBF, 2013) for internal surface vertical walls.

losses are reduced by about 88%. For the other colder cities, Milan and Zurich, the reductions are lower. The relative reductions can reach around 35% for Milan and 9% for Zurich. The north wall inner surface temperature under the climate of Barcelona is shown in Fig. 15 with and without the active closed water loop system. It is shown that the temperatures are significantly higher for the envelope with the active system than those of the envelope without the system. The same figures apply for Marseille. For Milan and Zurich, the difference is lower. 5.3. Parametric study

5.2. Base case simulation results The net heat gain or loss (meaning the absolute sum of the heat losses and heat gains during the simulation period) through the north wall are calculated for an envelope with the south-north active embedded pipe water loop system and for an envelope without this system and are shown in Fig. 14. Different solar irradiation control value (I_control) are investigated for the different cities. As shown in Fig. 14, the net heat losses through the north facade for the Mediterranean climates (Barcelona and Marseille) are drastically decreased. Also, we can even observe a heat gain for the city of Barcelona. The heat losses through the south facade (not shown) are slightly higher for the case with the active embedded-pipe system. They are increased by about 1–5% for the different cases compared to the reference with no active embedded-pipe system. As shown in Fig. 14 the proposed system shows high efficiency for the Mediterranean climates such as Barcelona or Marseille. For Barcelona, there is a net heat gain, meaning that the north wall heat losses are reduced by 113%. For Marseille, the north wall heat

The results of the previous section were for the base case. A lot of parameters affects the system’s performance and efficiency. In this section, we carry out a parametric analysis to identify the effect of some design or environmental parameters on the energy output of the active wall embedded-pipe water (fluid) loop system. 5.3.1. Effect of external surface solar absorptivity The effect of the external solar absorptivity is illustrated in Fig. 16 for the four cities. For each city, we have chosen a specific solar irradiation control value (I_solar) for each city, which are 250, 250, 300, and 400 W/(m2 K), for Barcelona, Marseille, Milan, and Zurich, respectively. Simulations are carried out for three solar absorptivity values: 0.6 (base case), 0.7, and 0.8. As expected, increasing the exterior surface solar absorptivity increases the system’s efficiency. As an illustration, the savings (relative to an envelope without the active loop system) for the city of Zurich increases from 4.5% (0.6 solar absorptivity) to 19.8% and 35.5% when increasing the solar absorptivity to 0.7 and 0.8, respectively. For the Mediterranean climates, this will increase the sys-

Fig. 14. The net heat gain/loss through the north wall for an envelope with the south-north active embedded pipe water loop system and for an envelope without this system for the cities of (a) Barcelona, (b) Marseille, (c) Milan, and (d) Zurich for different solar control value (I_control). (Note: the four sub-figures do not have the same y-axis scale).

M. Ibrahim et al. / Solar Energy 147 (2017) 22–36

Fig. 15. The north wall inner surface temperature for an envelope with and without the active loop system under the climate of Barcelona.

tem’s performance and heat gain to the inside. However, this increase in the exterior surface solar absorptivity may cause an overheating problem during the summer season.

5.3.2. Effect of water (fluid) mass flow rate Simulations are carried out for four mass flow rates: 0.0055 kg/s (20 kg/h), 0.011 kg/s (40 kg/h) (base case), 0.05 kg/s (180 kg/h), and 0.1 kg/s (360 kg/h) for the four climates. The effect of the fluid mass flow rate on the net heat losses/gains through the north façade is shown in Fig. 17. As illustrated in Fig. 17, there is an optimal mass flow rate which results in the minimum heat losses or maximum net heat gain. Among the four values tested, the optimal mass flow rate is 0.05 kg/s for all climates. Taking Milan as an example, the savings are around 25%, 37%, 43%, and 42% for the mass flow rates of 0.0055 kg/s, 0.011 kg/s, 0.05 kg/s, and 0.1 kg/s, respectively. However, it should be noted that the circulating pump’s power consumption will increase with the increase in the mass flow rate.

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5.3.3. Effect of indoor set-point temperature The indoor set-point temperature has a significant effect on the performance of the wall active loop system. The inner surface temperatures are highly influenced by the indoor air temperature acting as boundary condition. In all the previous simulations, the indoor air temperature is kept at 19 °C all time. To examine the influence of this parameter, different indoor temperature setpoint schedules are considered: 19 °C (all time), 18 °C (all time), and night set-back (NS) (19 °C from 7:00 till 23:00 and 16 °C from 24:00 till 6:00). The outside convective heat transfer coefficient is set to 17 W/(m2 K) and I_solar is set to 250, 250, 300, and 400 W/ (m2 K) for Barcelona, Marseille, Milan, and Zurich, respectively. The net heat loss/gain through the north façade is shown in Fig. 18 for the different indoor air temperature set-point schedules. Decreasing the indoor air temperature set-point increases the system’s efficiency as shown in Fig. 18. As illustration for the climate of Milan, the system decreases the heat losses by about 37% (7.6 kW h/m2 as absolute value) when considering 19 °C as the set-point value, by about 50% (9.4 kW h/m2) when considering 18 °C as the set-point value, and by about 47% (9.1 kW h/m2) when considering the night set-back control strategy. 5.3.4. Effect of outside convective heat transfer coefficient The effect of the outside convective heat transfer coefficient (hout) on the heat losses of the north facade is shown in Fig. 19 for an envelope with the active loop system and for an envelope without this system. It can be shown that the convective coefficient has a significant influence on the active loop system’s efficiency and performance. For the static insulation envelope, hout does not have a great influence on the heat losses due to the presence of the insulation layer (7 cm). So we can see that the heat losses for the 3 values of the convection coefficient are very close. However, the heat

Fig. 16. The net heat gain/loss through the north wall for an envelope with the south-north active embedded pipe water loop system and for an envelope without this system for the cities of (a) Barcelona, (b) Marseille, (c) Milan, and (d) Zurich for different exterior surface solar absorptivity. (Note: the four sub-figures do not have the same y-axis scale).

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M. Ibrahim et al. / Solar Energy 147 (2017) 22–36

Fig. 17. The net heat gain/loss through the north wall for an envelope with the south-north active embedded pipe water loop system and for an envelope without this system for the cities of (a) Barcelona, (b) Marseille, (c) Milan, and (d) Zurich for different fluid’s mass flow rate. (Note: the four sub-figures do not have the same y-axis scale).

losses/gains vary significantly for the dynamic envelope with the embedded water loop system. For Milan city, the reduction in the net heat losses is 37% when considering a hout of 17 W/(m2 K). The reduction increases to 63% when considering a hout of 12 W/(m2 K) and decreases to 16% when considering a hout of 25 W/(m2 K). For Zurich city, the reduction in the net heat losses is 4.5% when considering a hout of 17 W/(m2 K). The reduction increases to 23% when considering a hout of 12 W/(m2 K) and decreases to 9% when considering a hout of 25 W/(m2 K), meaning that the active system behavior is worse than the static envelope for this specific case.

6. Discussion By examining the results of the base case and the parametric studies, we can conclude that the wall active closed water loop system performs very well in the Mediterranean climates. Under such climates, the heat losses through the north wall can be significantly reduced and there is even a net heat gain through this facade. For the other climates, the savings are lower. The relative reductions in the heat losses can reach around 35% for Milan and 9% for Zurich. From another aspect, the thermal comfort is enhanced by limiting the radiative heat losses between the occupants and the cooler wall due to the increased north wall inner surface temperature. As a result, the occupants can attain the same thermal comfort levels but with lower indoor air temperature set-point value. In fact, a decrease in heating set-point of 1 °C decreases the heating energy consumption up to 13% (Palmer et al., 2012). This issue was not addressed in this study, but we will examine this effect in future

work by considering a control strategy based on the operative temperature or thermal comfort (such as the Fanger’s predicted mean vote (PMV) index), and not only on the indoor air temperature. The system’s performance is highly affected not only by the climate but also on several design and operation parameters. Among these parameters are the indoor air set-point temperature, the exterior solar absorptivity, and the fluid mass flow rate. The outside convective heat transfer coefficient has a significant influence on the system’s efficiency as shown previously in Fig. 19. Thus, more realistic simulations are needed taking into consideration the wind speed and direction to calculate the heat transfer convection coefficient at the outside surface. Also, other parameters can affect the performance such as the piping’s diameter, the pipe’s position in the south façade, the control strategy of the circulating pump, and the exterior thermal insulation thickness. Thus, we need an optimization algorithm that takes all these parameters into account and finds the best combination of all these parameters to optimize the system’s efficiency. Yet, another aspect that must be looked at is the initial cost of such systems as well as the operating cost of the circulating pump. So, a single-variable (the heat losses/gains and energy consumption) and a multi-variable (energy consumption and initial costs) is necessary to determine the feasibility of this active wall water-loop system in different climates.

7. Conclusion and outlook In this study, we presented a novel concept of the active embedded-pipe envelope systems. The system consists of an active

M. Ibrahim et al. / Solar Energy 147 (2017) 22–36

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Fig. 18. The net heat gain/loss through the north wall for an envelope with the south-north active embedded pipe water loop system and for an envelope without this system for the cities of (a) Barcelona, (b) Marseille, (c) Milan, and (d) Zurich for different indoor air temperature set-point values (NS = night set-back). (Note: the four sub-figures do not have the same y-axis scale).

Fig. 19. The net heat gain/loss through the north wall for an envelope with the south-north active embedded pipe water loop system and for an envelope without this system for the cities of (a) Barcelona, (b) Marseille, (c) Milan, and (d) Zurich for different outside convective heat transfer coefficients. (Note: the four sub-figures do not have the same y-axis scale).

closed water loop pipes embedded in the building exterior fabric to collect and utilize the solar energy gain on the surface of south wall to decrease or offset the heat loss through the north wall. An exper-

imental study was carried out to test the system’s performance and a numerical model was developed and validated against experimental measurements. Numerical simulations were done for a

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M. Ibrahim et al. / Solar Energy 147 (2017) 22–36

base case and a parametric study was carried out to examine the system’s performance for different design and operating parameters. The main conclusion is that the system performs very well in the Mediterranean climates (or similar climates) and to a lower extent in the cooler ones. Its performance is highly dependent on several design, climatic, and operating variables which should be optimized to have the best efficiency. Accordingly, as a future work, the following issues will be addressed: – provide a multi-variable optimization algorithm of the wall active south-north embedded water pipe system taking into account different constraints or objective functions (including the operative temperature and comfort constraints) – provide design recommendations for such systems for different climates – test more robust control strategies for the circulating pump

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