What Provides Heat That is Controlled and Continuous

Continuous Heating

In continuous heating calorimetric studies, when the temperature is increased linearly with time, the thermal effects of glass transition and devitrification may overlap, making the analysis of the results very difficult.

From: Pergamon Materials Series , 2007

Hybrid laser-arc welding of aerospace and other materials

J. Zhou , ... P.C. Wang , in Welding and Joining of Aerospace Materials, 2012

Inverse Bremsstrahlung (IB) absorption

With the continuous heating of the laser beam, the temperature of the metal vapour inside the keyhole can reach much higher than the metal evaporation temperature, resulting in strong ionisation, which produces keyhole plasma. The resulting plasma absorbs laser power by the effect of Inverse Bremsstrahlung (IB) absorption. Equations [4.25] and [4.26] define the IB absorption fraction of laser-beam energy in plasma by considering multiple- reflection effects ⁶² :

[4.25] ${\alpha }_{\mathit{IB},1}=1-exp\left(-{\displaystyle {\int }_0^{s_0}{k}_{\mathit{pl}}\mathit{ds}}\right)$

[4.26] ${\alpha }_{\mathit{IB},\mathit{mr}}=1-exp\left(-{\displaystyle {\int }_0^{s_m}{k}_{\mathit{pl}}\mathit{ds}}\right)$

Here, α_IB,1 is the absorption fraction in plasma owing to the original laser beam; α_IB,_mr is the absorption fraction owing to the reflected laser beam. ${\displaystyle {\int }_0^{s_0}{k}_{\mathit{pl}}}\mathit{ds}\mathrm{and}{\displaystyle {\int }_0^{s_m}{k}_{\mathit{pl}}}\mathit{ds}$ are, respectively, the optical thickness of the laser transportation path for the first incident and multiple reflections, and k_pl is the plasma absorption coefficient owing to inverse Bremsstrahlung absorption ⁶³ :

[4.27] $k_{\mathit{pl}}=\frac{n_e{n}_i{Z}^2{e}^{6}2\pi }{6\sqrt{3m{\epsilon }_0^3\mathit{ch}{\omega }^3{m}_e^2}}{\left(\frac{m_e}{2\mathit{\pi k}{T}_e}\right)}^{0.5}\left[1-exp\left(-\frac{\omega }{k{T}_e}\right)\right]\overline{g}$

where Z is the average ionic charge in the plasma; ω is the angular frequency of the laser radiation; ε ₀ is the dielectric constant; k is the Boltzmann's constant; n_e and n_i are particle densities of electrons and ions; h is Planck's constant; m_e is the electron mass; T_e is the excitation temperature; c is the speed of light; and ḡ is the quantum mechanical Gaunt factor. For the weakly ionised plasma in the keyhole, the Saha equation ⁶³ can be used to calculate the densities of plasma species:

[4.28] $\frac{n_e{n}_i}{n_0}=\frac{g_e{g}_i}{g_0}\frac{{\left(2\pi m_{e}k{T}_e\right)}^{1.5}}{h^3}exp\left(-\frac{E_i}{k{T}_e}\right)$

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Hybrid laser-arc welding in aerospace engineering

J. Zhou , ... P.C. Wang , in Welding and Joining of Aerospace Materials (Second Edition), 2021

5.2.6.1 Inverse bremsstrahlung (IB) absorption

With the continuous heating of the laser beam, the temperature of the metal vapor inside the keyhole can reach much higher than the metal evaporation temperature, resulting in strong ionization, which produces keyhole plasma. The resulting plasma absorbs laser power by the effect of Inverse Bremsstrahlung (IB) absorption. Eqs. (5.25) and (5.26) define the IB absorption fraction of laser beam energy in plasma by considering multiple reflection effects [74]:

(5.25) ${\alpha }_{\mathit{iB},1}=1-exp\left(-\underset{0}{\overset{s_0}{\int }}{k}_{\mathit{pl}}\mathit{ds}\right)$

(5.26) ${\alpha }_{\mathit{iB},\mathit{mr}}=1-exp\left(-\underset{0}{\overset{s_m}{\int }}{k}_{\mathit{pl}}\mathit{ds}\right)$

here, α _{iB, 1} is the absorption fraction in plasma due to the original laser beam; α _iB,mr is the absorption fraction due to the reflected laser beam. $\underset{0}{\overset{s_0}{\int }}{k}_{\mathit{pl}}\mathit{ds}$ and $\underset{0}{\overset{s_m}{\int }}{k}_{\mathit{pl}}\mathit{ds}$ are, respectively, the optical thickness of the laser transportation path for the first incident and multiple reflections, and k _pl is the plasma absorption coefficient due to inverse Bremsstrahlung absorption [75]:

(5.27) $k_{\mathit{pl}}=\frac{n_e{n}_i{Z}^2{e}^{6}2\pi }{6\sqrt{3}m{\varepsilon }_0^3\mathit{ch}{\omega }^3{m}_e^2}{\left(\frac{m_e}{2\pi {\mathit{kT}}_e}\right)}^{0.5}\left[1-exp\left(-\frac{\omega }{{\mathit{kT}}_e}\right)\right]\overline{g}$

where Z is the average ionic charge in the plasma, ωis the angular frequency of the laser radiation, ɛ ₀ is the dielectric constant, k is the Boltzmann's constant, n _e and n _i are particle densities of electrons and ions, h is Planck's constant, m _e is the electron mass, T _e is the excitation temperature, c is the speed of light, and $\overline{g}$ is the quantum mechanical Gaunt factor. For the weakly ionized plasma in the keyhole, the Saha equation [75] can be used to calculate the densities of plasma species:

(5.28) $\frac{n_e{n}_i}{n_0}=\frac{g_e{g}_i}{g_0}\frac{{\left(2\pi m_e{\mathit{kT}}_e\right)}^{1.5}}{h^3}exp\left(-\frac{E_i}{{\mathit{kT}}_e}\right)$

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Thermal Engineering of Steel Alloy Systems

K.D. Clarke , in Comprehensive Materials Processing, 2014

12.10.4.1 Carbon Steel

Figure 7 shows continuous-heating TTA diagrams for Ck 15 and Ck 45 steels (equivalent to AISI 1015 and 1045), each starting from ferrite–pearlite initial microstructures (17,18).

The TTA diagram shows that the initial transformation from ferrite–pearlite to austenite (Ac₁) occurs at much higher temperatures with increasing heating rate, increasing from approximately 730 °C to nearly 790 °C. At high enough temperatures, austenite can be stable without carbon (911 °C in Figure 1), but at these lower temperatures austenite requires dissolved carbon to be stable. On heating, therefore, austenite nucleates at the carbides, which supply carbon to the growing austenite grains. Carbon must also diffuse from the carbide through the austenite to the growth front in order to allow further austenite growth, resulting in diffusion-controlled growth kinetics. Thus, the carbide dissolution and carbon diffusion kinetics are critical to austenite formation kinetics. In the case of a ferrite–pearlite microstructure, nucleation occurs at pearlite grain boundaries (3). For the continuous-heating TTA shown in Figure 7, it is also important to recognize that the ferrite-plus-pearlite-plus-austenite temperature range for the 1045 steel also increases with the increasing heating rate, which, as is shown in this chapter, is an effect of a coarse initial microstructure. Finally, in order to achieve homogeneous austenite, with respect to carbon, even higher temperatures are required, which can be very important to the subsequent transformations that occur when cooling this austenite to room temperature. Homogenization of substitutional alloying element segregation, as discussed in this chapter, requires very long times.

The effects of carbon content are also evident in Figure 7, with the Ac₃ temperature for the lower carbon steel significantly higher than for the 1045 steel, and thus requiring significantly higher austenitization temperature (approximately 860 °C vs 770 °C) for slow heating rates. The recommended standard austenitizing temperatures for annealing or normalizing these grades are similar – respectively 885 and 915 °C for 1015, and 845 and 900 °C for 1045 (13) – but do show that the 1015 composition requires higher temperatures for austenitizing. However, when heating rates are increased much higher than 100 °C s^{− 1}, the transformation temperatures are nearly equal, due to the kinetics of carbide dissolution and carbon diffusion required to form austenite: for the 1015 microstructure, the carbon diffusion distance is greater, since there is more ferrite, whereas for the 1045 microstructure, the larger volume fraction of carbides that must dissolve is likely the primary process controlling the rate of full austenitization. It is also important to note that the temperature at which homogeneous (i.e., no carbon gradients) austenite occurs is similar at high heating rates. The rate of carbon diffusion, as an interstitial element, is fast relative to substitutional elements, but heating rates above 1000 °C s^{− 1} are fast enough so that it must be considered a rate-limiting process; see Figure 5 (21,23). The starting microstructure in Figure 7 also shows significantly spheroidized pearlite for the 1045 starting microstructure, which suggests that the dissolution kinetics may be sluggish relative to a starting microstructure of pearlite with fine lamellae. Several studies have been performed on the effect of microstructural scale on 1045 austenitization kinetics (21,28–30). One study (21,28) compared a hot-rolled microstructure with 30 vol.% ferrite and a larger pearlite colony size with a normalized microstructure with 40 vol.% ferrite and a smaller pearlite colony size. Results are shown in Figure 8. Although the apparent effect on the transformation temperatures is small, the change in as-quenched hardness is dramatic: the normalized starting microstructure responds to austenitization heat treatment much more quickly than the coarser hot-rolled microstructure. These results show that, for industrial nonequilibrium heating or cooling rate processing, the use of a normalizing treatment to refine grain size is beneficial to achieve a faster austenitization response, which requires shorter heat treatments and lower maximum temperatures.

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Developments in pulsed and continuous wave laser welding technologies

J. Zhou , H.L. Tsai , in Handbook of Laser Welding Technologies, 2013

Inverse Bremsstrahlung (IB) absorption

As discussed above, with the continuous heating from the laser beam, the metal vapor inside the keyhole can be heated much above its evaporation temperature, resulting in strong ionization. Thus, keyhole plasma is produced. The resulting keyhole plasma can absorb a lot of laser power by the effect of Inverse Bremsstrahlung (IB) absorption, thus affecting the energy transport to the liquid metal. Equations [5.14] and [5.15] define the IB absorption fraction of laser beam energy in plasma by considering multiple reflection effects: ¹⁵

[5.14a] ${\alpha }_{\mathit{iB},1}=1-exp\left(-{\displaystyle {\int }_0^{s_0}{k}_{\mathit{pl}}\mathit{ds}}\right)$

[5.14b] ${\alpha }_{\mathit{iB},\mathit{mr}}=1-exp\left(-{\displaystyle {\int }_0^{s_m}{k}_{\mathit{pl}}\mathit{ds}}\right)$

where, α _iB,1 is the absorption fraction in plasma due to the original laser beam; α _iB,mr is the absorption fraction due to the reflected laser beam. ${\displaystyle {\int }_0^{s_0}{k}_{\mathit{pl}}\mathit{ds}\mathrm{and}}{\displaystyle {\int }_0^{s_m}{k}_{\mathit{pl}}}\mathit{ds}$ are, respectively, the optical thickness of the laser transportation path for the first incident and multiple reflections, and k_pl is the plasma absorption coefficient due to Inverse Bremsstrahlung absorption: ¹⁶

[5.15] $k_{\mathit{pl}}=\frac{n_e{n}_i{Z}^2{e}^{6}2\pi }{6\sqrt{3}m{\epsilon }_0^3\mathit{ch}{\omega }^3{m}_e^2}{\left(\frac{m_e}{2\mathit{\pi k}{T}_e}\right)}^{0.5}\left[1-exp\left(-\frac{\omega }{k{T}_e}\right)\right]\overline{g}$

where Z is the average ionic charge in the plasma, ω is the angular frequency of the laser radiation, ε ₀ is the dielectric constant, k is the Boltzmann's constant, n_e and n_i are particle densities of electrons and ions, h is Planck's constant, m_e is the electron mass, T_e is the excitation temperature, c is the speed of light, and $\overline{g}$ is the quantum mechanical Gaunt factor. For the weakly ionized plasma in the keyhole, the Saha equation ¹⁶ can be used to calculate the densities of plasma species:

[5.16] $\frac{n_e{n}_i}{n_0}=\frac{g_e{g}_i}{g_0}\frac{{\left(2\pi m_{e}k{T}_e\right)}^{1.5}}{h^3}exp\left(-\frac{E_i}{k{T}_e}\right)$

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Optimising wind turbine design for operation in cold climates

L. Battisti , in Wind Energy Systems, 2011

14.7.4 Intermittent (cyclic) hot gas heating

The energy penalties for a turbine using continuous heating for icing protection are extremely large and in some cases may be prohibitive. An advantageous method ( Battisti et al., 2006; Battisti and Fedrizzi, 2007) for icing protection in wind turbines can be accomplished by cyclical de-icing. Some ice is permitted to form on the surfaces but then is removed periodically during relatively short, intensive applications of heat.

A water film between the surface and the ice is caused by the application of heat, and permits the removal of the ice by aerodynamic forces. Because the heating is intermittent, heat is supplied successively to relatively small surfaces areas, and a constant load on a heat source is thus maintained. The total heat input for cyclical de-icing, therefore, can be greatly reduced from that required for continuous heating. If the large heating requirement associated with continuous hot gas heating can be reduced by the use of a cyclical hot gas de-icing system, the wind turbine performance penalties could be decreased from those incurred with the continuous heating system.

Even though the average thermal power level required for the two cases is comparable, the heating period (t _{Heat ON}) for the de-icing practice is much shorter than for the anti-icing system (t _{Heat ON}  + t _{Heat OFF}). Contrary to aeronautical de-icing systems that employ ratios up to 10 (Yasilik, et al., 1992; Gray et al., 1952), low values of this ratio, that is 0.5:1.5, can be typical for WT. This is caused by the different thermal conductivities of the materials used to manufacture the blade wall. The effect relates to the Fourier number which expresses the ratio between the thermal conduction and the thermal inertia of a body. Bodies with large Fourier numbers, that is, with high conductivity λ_mat and small thermal inertia ρc _th (such as aircraft wings made of aluminium alloy), allow large heat fluxes to readily diffuse through the volume. On the other hand, large heat fluxes cannot diffuse quickly through thick walls made of composites (low Fourier number), resulting in a large increase of the 'warm surface' (inner-side) temperature, without any significant heat diffusion at the 'cold surface' (outer-side) in short times. Therefore, in the latter case, low thermal power densities should be preferred for long warming periods. Numerical simulations (Battisti and Fedrizzi, 2007) have confirmed the validity of this approach. Figure 14.25 shows the intermittency factor variation versus the warm air temperature and LWC, for cold air temperature equal to −   3   °C. The intermittency factor drops as the hot air temperature is increased, due to the reduction of the time required for the ice to melt. The chart shows an asymptotic trend as the temperature increases. Figure 14.25 also shows that, during de-icing operation and for warm-air temperatures more than about 140   °C, the maximum inner surface temperature rises above 100   °C. This figure could affect the structural integrity of the blade, and has been used as a maximum working temperature. A minimum warm-air temperature between 70 and 80   °C is required for the de-icing practice; at lower temperatures, the intermittency factor rapidly approaches the limit of 1. Figure 14.26 compares the anti-icing and de-icing systems in terms of energy consumption ratio for operation at an ambient temperature of −   3   °C. The energy consumption ratio is defined in Eq.14.20:

14.25. Intermittency factor vs warm air temperature at T_a,∞  =   −   3   °C.

14.26. Energy consumption ratio vs warm air temperature at T _a,∞  =   −   3   °C.

[14.20] $E_{\tau }=\frac{{\dot{q}}_{\mathrm{ave},\mathrm{de}-\mathrm{icing}}\cdot T_{\mathrm{heat}-\mathrm{on}}}{{\dot{q}}_{\mathrm{anti}-\mathrm{icing}}\cdot \left(T_{\mathrm{heat}-\mathrm{on}}+T_{\mathrm{heat}-\mathrm{off}}\right)}=\frac{{\dot{q}}_{\mathrm{ave},\mathrm{de}-\mathrm{icing}}}{{\dot{q}}_{\mathrm{anti}-\mathrm{icing}}}\cdot \tau ={\dot{q}}_{\mathrm{anti}-\mathrm{icing}}\cdot \xi \cdot \tau$

where ξ is a magnification factor representing the fraction of extra power required for de-icing, and τ is the intermittency factor given in equation[14.3]. The graphs show that this ratio is less than 40% in all the cases considered, when de-icing is adopted instead of anti-icing. All the three curves present a minimum at about 120   °C, meaning that the best compromise between employed heating power and de-icing heat-on time is obtained in this condition. Thus, this operational condition is preferred, if it is compatible with the air heater power installed and with the blade at safe-working temperature.

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Crystallisation of nickel–phosphorus (Ni–P) deposits with medium and low phosphorus content

W. Sha , ... K.G. Keong , in Electroless Copper and Nickel–Phosphorus Plating, 2011

Abstract:

The crystallisation process in electroless nickel phosphorus (Ni–P) platings during continuous heating is investigated using the X-ray diffraction (XRD) analysis and the differential scanning calorimetry (DSC). The X-ray line broadening technique can estimate the grain size and the microstrain with the aid of software to separate the reflections of crystalline nickel from the amorphous phase. In the as-deposited condition, platings with 5–8   wt% phosphorus are mixtures of amorphous and nanocrystalline materials. In all the platings after continuous heating, the grain size increases sharply when the temperature increases above 400 °C. The microstrain in the platings decreases with increasing temperature.

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Modern applications and current status of green nanotechnology in environmental industry

Manita Thakur , ... Deepak Pathania , in Green Functionalized Nanomaterials for Environmental Applications, 2022

Microwave pyrolysis of carbonaceous compounds

In this method, 1   g sucrose has been dissolved in 4   mL water and mixed in 20 mL of orthophosphoric acid with continuous heating in the microwave oven (at 100  W) for 3   min 40   s. Further, there was a change in color from colorless to brownish black. The above solution was allowed to stand at room temperature in order to cool the solution. After this, 50   mL distilled water (DW) has been added, and again the solution was allowed to stand for some time. Brownish-black precipitates were obtained, which were collected by centrifugation at 4000   rpm for 10   min. The carbon nanoparticles have been washed many times with double distilled water (DDW) and dried at 40°C in a vacuum oven (Chandra et al., 2011).

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Adaptability and key concepts of the hot in-place recycling technology

Banting W.P. Sze , Yifu Zhang , in Research and Application of Hot In-Place Recycling Technology for Asphalt Pavement, 2021

2.3.2 Technical prerequisites—intermittent radiant heating technology

Bitumen in asphalt mixture is not a good heat conductor. During the heating process of asphalt pavement, direct and continuous heating will certainly lead to the overheating, burning, or charring of asphalt mixture. According to the characteristics of asphalt materials, bitumen should not stay under the temperature over 180°C persistently. Otherwise, it will lose the binder and viscosity in the asphalt mixture. As a result, the asphalt mixture is unable to maintain the required mechanical properties.

During the short time for construction work, it is difficult to heat the asphalt mixture to the operating temperature (140°C–170°C) for pavement and meanwhile, allow the heat to permeate to a certain depth (4–6   cm). Only the intermittent radiant heating can attain the ideal temperature as well as the required permeation depth.

For the intermittent radiant heating system, the working principle is to ignite fuel on the surface of a special microporous ceramic material, in order to generate extremely powerful (high) thermal energy for pulse heating of road pavement, as shown in Fig. 2.4. The heating control system allows the upper and lower limits of temperature to be set according to the properties of asphalt mixture. High thermal energy is transiently generated to heat road pavement for a short period. When asphalt mixture attains the upper limit of temperature, heating stops at the preset moment. The interruption is a heat-retaining process, which allows thermal energy to permeate into the depth of road pavement. When the pavement surface temperature drops to a certain preset lower limit, the heating control system automatically activates and heating starts again, as shown in Fig. 2.5. After several repeated cycles, the asphalt mixture can attain to a temperature ideal for pulverization and paving.

Figure 2.4. Working principle of intermittent radiant heating technology.

Figure 2.5. Temperature control of heating system.

Liquefied petroleum gas is used as the fuel. The special ceramic material acts as a thermal radiator on the intermittent heating panel, as shown in Fig. 2.6. Without naked flame during the heating process, asphalt aging during heating can be minimized. Heating efficacy and energy utilization are high. To prevent heat loss, additional heating-retaining panels are mounted next to the heating panel to allow heat to permeate into sufficient depth, as shown in Fig. 2.7.

Figure 2.6. Heating panel working.

Figure 2.7. Heat-retaining panel for heat insulation.

According to the tests conducted by Xi'an Road Construction Testing Center (formerly Road Construction Testing Center of Ministry of Transport), after the pavement being heated by the heating wall for 15   minutes, the temperature at the depth of 80   mm from the road surface could reach to 104°C, and a maximum depth of 81   mm were softened. Annex B shows the report.

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Adsorption Refrigerators and Heat Pumps

R.E. CRITOPHDr., in Carbon Materials for Advanced Technologies, 1999

5.2.1 Multiple beds

In the basic continuous adsorption cycle illustrated previously in Fig. 5b two adsorbent beds are heated and cooled out of phase in order to provide continuous heating or cooling. It is possible to recover some of the adsorption heat rejected by each bed and use it to provide some of the heat required by the other bed. This might be done by the use of a circulating heat transfer fluid or a heat pipe. Meunier [ 5] first systematically looked at the potential gain in COP that might be obtained by such heat recovery, both as a function of the approach temperature of the beds donating and receiving heat and of the number of beds. The number of beds is not limited to two and the COP increases with the number of beds that it is possible to transfer heat between. There is of course a practical limitation, but it is possible to calculate the theoretical benefit of employing heat exchange between any number of beds.

Heat management within the cycle is best understood with reference to Fig. 11 showing the effective specific heat dQ/dT of one of the beds as a function of its temperature T as it completes a cycle. The top curve is the effective specific heat of a conventional bed during the heating phase. The heating phase is assumed to start with an isostere during which the bed increases in pressure from that of the evaporator to that of the condenser. Since there is no desorption, the effective specific heat of adsorbent plus adsorbate per unit mass of adsorbent (dQ/dT) is constant. The second part of the curve shows the effect of desorption occurring along an isobar. The pressure is maintained constant by the ideal condenser which rejects heat at a constant temperature. dQ/dT is high at lower temperatures both because there is more adsorbate which imposes a sensible heat load and because the concentration change per unit temperature is higher giving a larger desorption heat load. As the temperature rises and the concentration falls dQ/dT will tend to the specific heat of the carbon.

Fig. 11. Effective specific heat v. Temperature in a two-bed regenerative cycle

The lower curve shows the effective specific heat during adsorption. Since the process is one of cooling, the curve is negative. Meunier [5] showed how the two curves can be used to determine how much rejected heat from one bed can be used to heat another. The two rules guiding the calculation are: that the area under the curves between two temperature limits is equal to the total heat entering or leaving between those temperatures, and that heat can only flow from high to lower temperatures. In the two bed case illustrated, the heat rejected from the adsorbing bed between T₃ and T₆ , equal to Q_reg may be transferred across temperature difference ΔT to the adsorbing bed, which it heats from T₁ to T₅ . This reduces the total heat input from Q_in + Q_reg to Q_in with a consequent improvement in COP. There is also a reduction in the rejected heat to Q_reject . A two-bed cycle with heat recovery across a ΔT of 10°C can increase the refrigeration COP by 50% or more as shown in Fig. 12.

Fig. 12. Improvement in refrigeration COP possible with an idealised two-bed cycle

Greater numbers of beds allow more regeneration of heat but the benefit of increasing the number of beds drops off rapidly. Table 2 below taken from Meunier [5] shows how the COP changes with number of beds in a particular case.

Table 2. Variation in cooling COP of a zeolite – water regenerative cycle with evaporating, condensing, adsorption rejection and maximum desorption temperatures 0°C, 40°C, 50°C and 350°C respectively.

Number of beds	Cooling COP
1	0.425
2	0.684
4	1.008
6	1.293
∞	1.852

Considering the extreme case when there are an infinite number of beds (Fig. 13) and ideal heat transfer, the maximum amount of recovered heat is calculated by reflecting the lower curve in the upper half of the diagram, as is shown by the dashed line. The input heat that must be obtained from an external source is shown by the shaded area. The COP is the maximum that may be obtained from a single effect adsorption machine.

Fig. 13. Heat management with an infinite number of beds

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Traditional Houses and Energy Use in Barra and South Uist

Peter Whyman , in Energy for Rural and Island Communities: Proceedings of the Conference, Held at Inverness, Scotland, 22–24 September 1980, 1981

ABSTRACT

The Hebridean black house had at its centre an open peat fire which burned continuously. The traditional "Department plan" three apartment house was designed for continuous heating by two open fireplaces, one at either gable.

The difficulty of obtaining adequate supplies of peat and the advent of electricity have brought about the introduction of an intermittent heating cycle which is costly and inefficient in modifying the damp interior climate.

Re-introduction of continuous heating must depend on two factors: the cheapest possible fuel and the most economical use of fuel.

Simple mechanical aids to the extraction and drying of peat could make the use of this fuel in improved continuous burning appliances possible at low cost.

It is suggested that the black house was more efficient in its use of energy than the "Department plan" house and that given a clean source of heat it offers a better model for development.

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