Open access peer-reviewed chapter

Energy Storage in Concrete Bed

By Adeyanju Anthony Ademola

Submitted: February 5th 2018Reviewed: April 12th 2018Published: November 14th 2018

DOI: 10.5772/intechopen.77205

Downloaded: 205

Abstract

The energy storage ability and temperature arrangement of a concrete bed which was charged and discharged at the same time was examined mathematically in this research. This was carried out by modeling a single globe-shaped concrete which was utilized to simulate a series of points along the concrete bed axis. Charging and discharging mode of the system were compared for 0.0094, 0.013, and 0.019 m3/s air flow rates. Higher change in temperature response was detected between the charging and fluid to solid heat transfer process at the inception of the concrete bed and the heat gain by the cool air flowing inside the copper tube was fairly high. The analysis of energy storage efficiency was also carried out and it was noticed that the globe-shaped concrete of 0.11 m diameter has the highest storage efficiency of 60.5% at 0.013 m3/s airflow rate.

Keywords

  • energy storage
  • heat transfer model
  • globe-shaped concrete
  • charging and discharging

1. Introduction

Heat can be transferred in a concrete bed through the following means: (i) heat transfer by convection from the bed wall to the fluid flowing inside the bed; (ii) heat transfer by convection from the globe-shaped concretes to fluid flowing in the bed, and this is known as fluid to particle mode; (iii) heat transfer by conduction from the bed walls to the globe-shaped concretes; (iv) heat transfer by conduction from one globe-shaped concrete to another, and this is known as particle to particle mode; (v) heat transfer by radiation; and (vi) heat transfer through fluid mixture [1]. The modes are as shown in Figure 1.

Figure 1.

Schematic showing the modes of heat transfers in concrete bed.

The particle to particle conduction mode can be further analyzed in axial and radial directions. Heat transfer through radiation mode will actually be important at higher temperatures.

Practically, it is found that two or three of the modes discussed earlier on can occur simultaneously. For example, the conduction between the particles may be affected by the convection between the particles and the fluid. This interaction among the different modes is one of the main reasons for the difficulty in correlating the total heat transfer and analyzing the experimental data in this field [2].

This research carried out numerically the temperature distribution in a concrete bed and also looked into the ability of a concrete bed to store energy simultaneously.

2. Review of literature

Anzelius [3] is the first author to publish a paper in heat transfer through packed beds but Schumann [4] is usually the first reference cited in most literature [1]. Both authors made some assumptions in order to find solution to equations that guide heat transfer for an incompressible fluids passing uniformly through a bed of solid particles with perfect conductivity. The following are the heat transfer equations derived for the system:

TsTs,0TsTs,0=1eYZΣYnMnyz=eYZΣZnMnE1
TfTsTf,0Ts,0=1eYZΣYnMnyz=eYZΣZnMnyzE2

where Ts is the solid temperature, Tf is the temperature of fluid and Y and Z are dimensionless quantities. Eqs. (1) and (2) was solved simultaneously and the result were analyzed graphically to form Schumann curves. Knowing the outlet air and bed temperature, these curves can be used to determine the volumetric heat transfer coefficients and also the heat transfer coefficients of a packed bed undergoing heat exchange with a fluid provided the following conditions which were the simplifying assumptions made by Schumann were satisfied:

  1. Due to the infinitesimal nature of the solid particles, resistance to heat transfer was so small.

  2. Resistance to conduction heat transfer in the fluid was so small.

  3. At any section of the bed, the heat transfer rate from fluid to solid or from solid to fluid was directly proportional to their average difference in temperature between the solid and fluid within the bed.

  4. The transport properties of solid and fluid were not dependent on temperature, for example the density.

Furnas [5] utilized and expanded the Schumann curves to cover more range of temperatures and suggested an empirical relation for determination of heat transfer coefficient as shown in Eq. (3):

hv=BG0.7T0.3101.68ε3.56ε2dp0.9E3

where, hv is the volumetric heat transfer coefficient. B is a constant dependent on the bed material, G is the mass velocity of the fluid, T is the average air temperature, dp is the particle diameter and ε is the porosity.

Saunders and Ford [6] utilized dimensional analysis to derive correlations to calculate heat transfer coefficient. The research was for spherical shaped alone and the application is not suitable for other solid particle geometries.

Another correlation for the determination of heat transfer coefficient between gases and randomly packed solid spheres was postulated by Kays and London [7]. Using the Colburn j-factor, the relationship was written as:

jh=0.23Rep0.3E4
where,jh=St.Pr2/3
200<Rep<50,000and0.37<ε<0.39

Löf and Hawley [8] studied heat transfer between air and packed bed of granitic gravel. Unsteady state heat transfer coefficients were correlated with the air mass velocity and particle diameter to obtain the equation:

h=0.652G/dp0.7E5

This was calculated for 8 mm < dp < 33 mm; 50 < Rep 500 and temperature range of 311–394 K. The author reached a conclusion that the temperature of the entering air had no appreciable effect on the hat transfer coefficient.

Leva et al. [9] studied and analyzed heat transfer coefficient between smooth spheres of low thermal conductivities and fluids (air and carbon dioxide) in packed beds and tubes of 50.8 and 6.4 mm diameters, respectively. The ratio of particles to tube diameters was varied from 0.08 to 0.27; gas flow rate was of Reynolds number range 250 to 3000. Correlation of film coefficient was determined as:

h=3.50kDte4.6DpDtDpGµ0.7E6a

which is approximately:

h=0.40k/DtDpG/µ0.7E6b
Or,N=0.4Re0.7E6c

Maximum film coefficient was predicted and verified at a value of Dp/Dt equal 0.153.

Riaz [10] and Jefferson [11] studied the dynamic behavior of beds undergoing heat exchange with air using single and two phased modes. By incorporating factors of axial bed conduction and intra-particle resistance, which Schumann ignored, the heat transfer coefficients were evaluated and found to be 1 + Bi/5 times smaller than those predicted using Schumann curves.

Ball [12], Norton [13], Meek [14], Bradshaw and Meyers [15], Harker and Martyn [16] and also, Bouguettaia and Harker [17] have all researched on various packed beds using air and other gases as fluids and have developed correlations involving the heat transfer coefficient.

3. Methodology

3.1. Heat transfer model for a globe-shaped concrete bed

The modeling of heat transfer in a concrete bed was carried out mathematically. It was done through a single globe-shaped concrete which was simulated mathematically to represent series of points along the concrete bed axis.

A one dimensional finite difference formulation was used in modeling the single globe-shaped concrete material, where heat conduction to neighboring globe-shaped concrete was ignored.

Using this assumption reduced the globe-shaped concrete model to that of an isolated sphere in cross flow, where the total surface area of the sphere was exposed to convection. Also, the thermal properties of the materials within the bed accounted for temperature dependence.

3.2. Finite difference formulation of a single spherical shaped concrete material

Since conduction to other globe-shaped concrete has been neglected, the geometry allows the concrete to be reduced to one dimension along its radius.

A finite difference method was utilized to model this mathematically, [18]. For this approach, the globe-shaped concrete can be characterized by three different nodal equations:

  1. a general, interior node

  2. the center node

  3. the surface node

All exposed to convection as shown in Figure 2.

Figure 2.

Schematic showing the cross-section of the globe-shaped concrete materials imbedded with copper tube.

For the general and interior node within the globe-shaped concrete model, the conduction equation for T(r,t) is:

ρcCcTt=1r2rKcr2Tr+q̇rtE7

where Cc = specific heat of concrete.

Kc=thermal conductivity of concrete.

q̇=heat generation.

And this equation was represented in finite difference form.

The specific heat, thermal conductivity, and the heat generation, are temperature dependent and varied with the temperature along the radial direction.

Because the thermal properties are functions of temperature, and consequently functions of the globe-shaped concrete radius, the finite difference equations are derived by the volume integration over a finite difference node.

Multiplying Eq. (7) by r2 and integrating both sides of the equation from rn – Δr/2 to rn + Δr/2 resulted to:

ρcCcTtrnΔr2rn+Δr2r2dr=rnΔr2rn+Δr2rKcr2Trdr+rnΔr2rn+Δr2r2q̇rtdrE8

The specific heat was assumed constant with respect to r, and therefore brought outside the integral.

By evaluating the integrals in Eq. (8) and representing the derivatives in finite difference form using the fully implicit method gives:

ρcCcTnZ+1TnZΔtrn+3rn33=rnΔr2rn+Δr22KcrdT+q̇rnΔr2rn+Δr2r2drE9
ρcCcTnZ+1TnZΔtrn+3rn33=2Kcr22dTΔrrnΔr2rn+Δr2+q̇nr33rnΔr2rn+Δr2E10
ρcCcTnZ+1TnZΔtrn+3rn33=Kcn+rn+2Tn+1Z+1TnZ+1ΔrKcnrn2TnZ+1Tn1Z+1Δr+qnrn+3rn33rnΔr2rn+Δr2E11

where TnZand TnZ+1indicate temperatures for an arbitrary node at times tZand tZ+1respectively.

Also,

Wn=WatTnZ+1E12

and,

Kn+=KTn+1Z+1+K2,atTnZ+1E13

also,

rn+=rn+Δr2E14
rn=rnΔr2E15

Eq. (11) can be rearranged and solved for TnZplus a known quantity which resulted to:

ρcCcTnZ+1TnZΔt=3Kcn+rn+2rn+3rn3Tn+1Z+1TnZ+1Δr3Kcnrn2rn+3rn3TnZ+1Tn1Z+1Δr+q̇n3rn+3rn33rn+3rn3E16
ρcCcTnZ+1ΔtρcCcTnZΔt=3Kcn+rn+2rn+3rn3Tn+1Z+1TnZ+1Δr3Kcnrn2rn+3rn3TnZ+1Tn1Z+1Δr+q̇nE17

Multiply Eq. (17) by Δt and divide by ρcCc resulted to:

TnZ+1TnZ=3Kcn+rn+2ΔtρcmCcmΔrrn+3rn3Tn+1Z+1TnZ+13Kcnrn2ΔtρcmCcmΔrrn+3rn3TnZ+1Tn1Z+1+3Δtq̇nρcmCcmE18
TnZ+3Δtq̇nρcmCcm=TnZ+13Kcn+rn+2ΔtρcmCcmΔrrn+3rn3Tn+1Z+1+3Kcn+rn+2ΔtρcmCcmΔrrn+3rn3TnZ+1+3Kcnrn2ΔtρcmCcmΔrrn+3rn3TnZ+1+3Kcnrn2ΔtρcmCcmΔrrn+3rn3Tn1Z+1E19

Collecting the like terms from Eq. (19) yielded:

TnZ+3Δtq̇nρcmCcm=3ΔtρcmCcmΔrKcnrn2rn+3rn3Tn1Z+1+1+3ΔtρcmCcmΔrKcnrn2+Kcn+rn+2rn+3rn3TnZ+13ΔtρcmCcmΔrKcn+rn+2rn+3rn3Tn+1Z+1E20

This resulting equation is valid for any general, interior node within the globe-shaped concrete 0 < rn < R.

At the center node, where rn = 0 the temperature profile is axisymmetric, and Tr=0,when r = 0 thus, the temperature on either side of the node is equal.

Tn1Z+1=Tn+1Z+1
likewise,Kcn=Kcn+

Eq. (20) simplified to:

TnZ+Δtq̇nρcmCcm=TnZ+16ΔtρcmCcmΔr2Kcn+Tn+1Z+1E21

This occur at rn = 0.

This simplified form of Eq. (20) was used to represent the center node.

The conduction through the surface of the globe-shaped concrete is equal to the convection at the surface.

KTrat,r=R=UCTr=RTE22

However, this boundary condition cannot be directly represented in finite difference form, since such formulation requires a volume element and Eq. (22) applies at a point.

Instead a first law energy balance was utilized to obtain the nodal equation for the surface of the globe-shaped concrete. This energy balance can be written as:

ĖinĖout+Ėgen=ĖstE23

where,

Ėin=KATrE24
Ėout=UCATTgE25
Ėgen=q̇VE26
Ėst=ρCVTtE27

Representing Eq. (23) in a finite difference form consistent with Eq. (20) and (21) resulted to:

KATrUCATTg+q̇V=ρCVTtE28

This can be written in finite difference form to give:

4πrn2KnTnZ+1Tn1Z+1Δr4πrn2UCTnZ+1T+43πrn3rn3q̇n=43πρCnrn3rn3TnZ+1TnZΔtE29

where, Uc = convection coefficient.

Solving for TnZplus known quantities involving q̇and Uc, in a similar manner to Eq. (20) and (24) resulted to:

4πrn2KnTnZ+1Δr+4πrn2KnTn1Z+1Δr4πrn2UCTnZ+14πrn2UCT+43πrn3rn3q̇n=43πρCnrn3rn3ΔtTnZ+143πρCnrn3rn3ΔtTnZE30

Multiply Eq. (30) by Δt and divide by 43πρrn3rn3resulted to:

TnZTnZ+1=3ΔtΔCnrn2Knrn3rn3TnZ+13ΔtΔCnrn2rn3rn3Tn1Z+1+3ΔtUCρCnrn2rn3rn3TnZ+1+3ΔtUCρCnrn2rn3rn3Tq̇nΔtρCnE31
TnZ+ΔtρCnq̇n+3ΔtUCρCnrn2rn3rn3=3ΔtΔCnrn2Knrn3rn3Tn1Z+1+1+3ΔtρCnΔrrn2Knrn3rn3+3ΔtUCρCnrn2rn3rn3TnZ+1E32
TnZ=3ΔtΔCnrn2Knrn3rn3Tn1Z+1+1+3ΔtρCnΔrrn2Knrn3rn3+3ΔtUCρCnrn2rn3rn3TnZ+1ΔtρCnq̇n3ΔtUCρCnrn2rn3rn3E33

Eqs. (20), (21) (32) and (33) constitute a system of algebraic equations for heat transfer modeling in globe-shaped concrete.

4. Result and discussion

The values of Eq. (33) are obtained from the values in Table 1. Since the thermal properties are constant, average temperatures could therefore be used to determine thermal properties of bed materials.

ParametersValues
Airflow rate0.01316 m3/s (28 cfm)
Air—density1.07154 Kg/m3
Air—specific heat capacity1008 J/Kg K
Concrete—density2400 Kg/m3
Concrete—specific heat capacity1130 J/Kg K
Copper tube—density8900 Kg/m3
Copper tube—specific heat capacity384 J/Kg K
Area of globe-shaped concrete0.013 m2
Area of copper tube + header0.664 m2
Volumetric heat transfer coefficient106.5 W/m3 K

Table 1.

Modeling parameters.

The following data were obtained from the theoretical/mathematical modeling carried out on thermal performance of packed bed energy storage system as shown in Figure 3.

Figure 3.

Schematic of the storage tank systems.

The following are the definitions of the symbols:

Time = the interval time of measurements, in minutes.

Ts-in = the inlet air temperature to the packed bed storage tank in °C.

Ts-out = the outlet air temperature from the packed bed storage tank in °C.

Tt-in = the inlet air temperature to the copper tube in °C.

Tt-out = the outlet air temperature from the copper tube in °C.

TA1, TA2, TA3, and TA4 = the air stream temperatures (°C) through the bed at different heights of the storage tank 117.5, 235, 352.5, and 470 cm, respectively.

Tci1, Tci2, Tci3, and Tci4 = the core temperatures of the globe-shaped concrete (°C) through the bed at different heights of the storage tank 117.5, 235, 352.5, and 470 cm, respectively.

Tti1, Tti2, Tti3, and Tti4 = the temperatures of air flowing inside the copper tube (°C) through the bed at different heights of the storage tank 117.5, 235, 352.5, and 470 cm, respectively.

Tct1, Tct2, Tct3, and Tct4 = the temperatures of the contact made between globe-shaped concrete and imbedded copper tube (°C) through the bed at different heights of the storage tank 117.5, 235, 352.5, and 470 cm, respectively.

Tt1, Tt2, Tt3, and Tt4 = the surface temperatures of the copper tube (°C) through the bed at different heights of the storage tank 117.5, 235, 352.5, and 470 cm, respectively.

The results of the experimentation were shown in Figures 4, 5, 6 for globe-shaped concrete of size 0.11; 0.08 and 0.065 m diameter respectively while the discharging only temperature measurements were shown in Figures 79 respectively for air flow rate of 0.0094, 0.013, and 0.019 m3/s.

Figure 4.

Average temperature measurement of charging packed bed storage system for globe-shaped concrete of size 0.11 m diameter and flow rate of 0.0094, 0.013, and 0.019 m3/s.

Figure 5.

Average temperature measurement of charging packed bed storage system for globe-shaped concrete of size 0.08 m diameter and flow rate of 0.0094, 0.013, and 0.019 m3/s.

Figure 6.

Average temperature measurement of charging packed bed storage system for globe-shaped concrete of size 0.065 m diameter and flow rate of 0.0094, 0.013, and 0.019 m3/s.

Figure 7.

Average temperature measurement of discharging packed bed storage system for globe-shaped concrete of size 0.11 m diameter and flow rate of 0.0094, 0.013, and 0.019 m3/s.

Figure 8.

Average temperature measurement of discharging packed bed storage system for globe-shaped concrete of size 0.08 m diameter and flow rate of 0.0094, 0.013, and 0.019 m3/s.

Figure 9.

Average temperature measurement of discharging packed bed storage system for globe-shaped concrete of size 0.065 m diameter and flow rate of 0.0094, 0.013, and 0.019 m3/s.

Figure 10 presents the comparison of the temperature variations with time at Ts-in, Ts-out, Tt-in, Tt-out, TA1, TA2, TA3, TA4, Tci1, Tci2, Tci3, Tci4, Tti1, Tti2, Tti3, Tti4, Tct1, Tct2, Tct3, Tct4, Tt1, Tt2, Tt3, and Tt4 during the simultaneous charging and discharging while Figure 11 presents for discharging only. The comparisons were presented for air flow rates of 0.0094, 0.013, and 0.019 m3/s.

Figure 10.

Comparison of average temperature measurement of charging packed bed storage system for globe-shaped concrete of size 0.065, 0.08, 0.11 m in diameter and flow rate of 0.0094, 0.013, and 0.019 m3/s.

Figure 11.

Comparison of average temperature measurement of discharging packed bed storage system for globe-shaped concrete of size 0.065, 0.08, 0.11 m in diameter and flow rate of 0.0094, 0.013, and 0.019 m3/s.

These figures show that the difference of the temperature response between the charging and fluid to solid heat transfer process at the initial period (<30 min) of the packed bed was large (large inlet–outlet temperature difference means large heat supply), and the heat recovered by the cool air (approximately 27°C) flowing inside the copper tube was fairly high (larger inlet–outlet temperature difference compared with the later period indicates larger heat recovery).

Therefore, a relatively large part of the heat supplied by the simulated air heater was used to heat the air flowing inside the copper tube through conduction and convection and also stores the rest for continuous usage.

The following are the storage efficiency for globe-shaped concrete of size 0.11 m, 0.08 m and 0.065 m diameter at airflow rate of 0.0094, 0.013 and 0.019 m3/s (Figure 12):

Figure 12.

Storage efficiency of simultaneous charging and discharging packed bed storage system for globe-shaped concrete of diameter 0.065, 0.08, 0.11 m and air flow rate 0.0094, 0.013, and 0.019 m3/s.

For 0.11 m diameter globe-shaped concrete:

  1. Storage efficiency at air flow rates of 0.0094 m3/s = 40.7%

  2. Storage efficiency at air flow rates of 0.013 m3/s = 60.5%

  3. Storage efficiency at air flow rates of 0.019 m3/s = 57.5%

For 0.08 m diameter globe-shaped concrete:

  1. Storage efficiency at air flow rates of 0.0094 m3/s = 23.5%

  2. Storage efficiency at air flow rates of 0.013 m3/s = 51.3%

  3. Storage efficiency at air flow rates of 0.019 m3/s = 50.2%

For 0.065 m diameter globe-shaped concrete:

  1. Storage efficiency at air flow rates of 0.0094 m3/s = 14.8%

  2. Storage efficiency at air flow rates of 0.013 m3/s = 35.06%

  3. Storage efficiency at air flow rates of 0.019 m3/s = 40.3%

5. Conclusion

The study led to the following findings and conclusions:

  1. The mathematical model developed can accurately predict the temperature within the concrete bed for energy storage purpose.

  2. The steady intermittent input temperature variation actually led to continuous discharge temperature at the copper tube outlet.

  3. The mathematical model may be extended to specify the packed bed storage system dimensions.

  4. Globe-shaped concrete of 0.11 m diameter has the highest storage efficiency of 60.5% at 0.013 m3/s airflow rate.

© 2018 The Author(s). Licensee IntechOpen. This chapter is distributed under the terms of the Creative Commons Attribution 3.0 License, which permits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited.

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Adeyanju Anthony Ademola (November 14th 2018). Energy Storage in Concrete Bed, HVAC System, Mohsen Sheikholeslami Kandelousi, IntechOpen, DOI: 10.5772/intechopen.77205. Available from:

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