High-efficiency electrochemical thermal energy harvester using carbon nanotube aerogel sheet electrodes

Conversion of low-grade waste heat into electricity is an important energy harvesting strategy. However, abundant heat from these low-grade thermal streams cannot be harvested readily because of the absence of efficient, inexpensive devices that can convert the waste heat into electricity. Here we fabricate carbon nanotube aerogel-based thermo-electrochemical cells, which are potentially low-cost and relatively high-efficiency materials for this application. When normalized to the cell cross-sectional area, a maximum power output of 6.6 W m−2 is obtained for a 51 °C inter-electrode temperature difference, with a Carnot-relative efficiency of 3.95%. The importance of electrode purity, engineered porosity and catalytic surfaces in enhancing the thermocell performance is demonstrated.


Supplementary Tables
Supplementary Table 1 Inter-electrode spacing (m) 0.025 Internal resistance of the thermocell from a E-I curve shown in For the used electrolyte flow of 6.6 × 10 -6 m 3 s -1 , laminar flow of electrolyte is expected in a low Reynolds number regime (Re ~ 650). Using the above equation, the calculated Sh number is ~139, and correspondingly, the theoretically limiting mass transfer coefficient for a flat plate in this electrolyte is predicted to be 5.76 × 10 -6 m s -1 . While the measured mass transport number of the CNT aerogel electrode (5.19 × 10 -6 m s -1 ) is nearly as high, the measured mass transport number of the CNT buckypaper electrode is much lower (2.51 × 10 -6 m s -1 ).

Supplementary Note 2: Calculation of energy conversion efficiency of thermocells
The energy conversion efficiency (η) of a thermocell is defined as the ratio of maximum electrical power output (P max ) from the cell to thermal power flowing through the cell: (1) where V oc and I sc are the open-circuit voltage and the short-circuit current, respectively, κ is the thermal conductivity of electrolyte, A c is th cross s ctional ar a of th c ll, ΔT is the absolute temperature difference between two electrodes, and d is the inter-electrode spacing.
Applying the relationships, V oc = αΔT and I sc = αΔT/R cell , to Eq. (1), where α is the electrochemical Seebeck coefficient, R cell is the internal resistance of the thermocell, leads to the following equation: (2) The theoretical efficiency relative to Carnot efficiency (η r ) can be expressed by the following equation, when Eq. (2) is divided by the Carnot efficiency (η c = ΔΤ/T H ): where σ eff represents the effective conductivity in thermocell, analogous to the electrical conductivity in thermoelectrics. We note that the effective conductivity is evaluated from the internal resistance of the cell.
The resulting calculated Carnot-relative efficiency was = 3.95% for the optimized cylindrical thermocell. The parameters used in the calculation are summarized in Supplementary Table 3.

Supplementary Note 3: Electrochemical Seebeck coefficient and the effective ionic conductivity
A 0.4 M potassium ferri/ferrocyanide aqueous solution with a concentration close to saturation was used as the thermoelectric electrolyte. The electrolyte was prepared using deionized (DI) water and degassed prior to use by bath sonication. The freshly prepared electrolytes were used immediately to avoid the effects of l ctrolyt d gradation. 50 μm thick CNT sh ts with id ntical ar a of 1.0 cm 2 were used as electrodes.
Each CNT sheet electrode was connected to a 0.5 mm diameter platinum (Pt) wire using silver paste which was used to minimize the contact resistance. The contact was then covered by insulating paint to prevent possible artifacts due to interaction between the silver paste and the electrolyte.
A U-shaped cell equipped with liquid flowing pocket at each side was utilized for the measurement of electrochemical Seebeck coefficient (see Supplementary Fig. 5a). The distance between the two half-cells is 3 cm and the temperature of each side cell was controlled by circulating water from a thermostatic bath with an accuracy of ± 0.1 °C. Electrode temperatures were measured using thermocouple probes that were placed in close proximity to the electrode for each half-cell.
The thermoelectric coefficient of the redox couple was obtained by measuring the temperature dependence of the potential difference over a temperature range from 0 to 20 o C with an increment of ±2 o C.
The potential and current output from the cell was measured using a voltage-current meter (Keithley 2000 multimeter) with 0.002% DC voltage accuracy from 100 nV to 1 KV. As shown in Supplementary Fig. 5b, the thermoelectric coefficient was measured to be ~1.43 mV K -1 , which is in good agreement with previous reports.
In thermoelectric devices, an electrical conductivity of the thermoelectric is used to calculate the energy conversion efficiency. It implies that an electrical potential gradient ( V) is the dominant driving force to transport charges (electrons or holes) in the thermoelectric. However, mass transport (i.e., ion conduction) in thermocells results from both the diffusion processes based on electrical potential gradient ( V), thermal gradient (Soret diffusion, T), and concentration gradient (Fickian diffusion, c), and the convective process based on density gradient ( p). In other words, the ion conduction in thermocell is forced by the sum of the above driving forces, not solely by an electrical driving force. Moreover, the discharge behavior of the thermocell is determined by three primary internal resistances (i.e., activation, ohmic and mass transport overpotentials). Therefore, we cannot simply plug an ion conductivity into Eq. (3) but the effective conductivity ( ) should be evaluated from the internal resistance of the cell, i.e., the slope of E-I curve, at a given geometry of thermocell. For instance, the internal resistance of the cylindrical thermocell is measured as ~29 Ω from th E-I curve shown in Fig. 5d. With the cross sectional area of the cell (7.1 × 10 -6 m 2 ) and the inter-electrode spacing (0.025 m), the effective ionic conductivity ( )) is calculated to be ~1210 mS cm -1 .

Supplementary Note 4: Thermal transport in the cylindrical thermocell
Thermal transport in thermocells is generated not only by the heat flow due to thermal conduction through the electrolyte, but also by the additional flow due to all convective processes. Convective heat transfer can be driven by the temperature difference when the electrodes are held in a certain configuration (e.g., the cold above the hot electrode) and by the difference between the densities of the reactants and products in the ongoing reactions at the hot and the cold electrodes. In order to understand how the thermal transport is generated in the present cylindrical thermocell, we conduct a theoretical analysis on the thermal transport as follows: For an analysis of a quenched convection heat transfer in a horizontally long enclosure between the hot and cold ends, which represents the thermocell of a cylindrical enclosure, consider a two-dimensional enclosure of height H and horizontal length L, with infinite depth, as shown in Supplementary Fig. 6.
In the internal natural convection, the Rayleigh number based on the enclosure height is defined as (4) Here, g is the gravity, and are the fluid densities near heated and cooled walls, respectively, H is the height of the two-dimensional rectangular enclosure which is configured with two infinite horizontal walls, and and are the thermal diffusivity and kinematic viscosity of fluid, respectively.
For the natural convection condition, the heat transfer rate between the hot and cold end walls is given by 2 : In the case of shallow enclosure ( ), the above convective heat flow from the hot end wall can be diffused vertically downward from the warm upper branch of the circulation flow to the lower branch before reaching the cold end wall (see the dashed arrow in Supplementary Fig. 6). In other words, the convective heat transfer will be quenched down in the middle of the enclosure and diffused back to the hot end wall area. The vertical heat diffusion rate is given by 2 : If the vertical diffusion rate of Eq. (6) is higher than the convective heat transfer rate of Eq. (5), the energy carried by the upper stream cannot reach the cold end. In this case, the two branches of horizontal counter-flows diffuse to make good thermal contact, which diminishes the convection flow and thus, results in conduction heat transfer dominating across the electrolyte solution, i.e., (7b) In applying the above criterion to the present cylindrical enclosure, Eq. (4) should be modified for the cylindrical enclosure using the characteristic length of the infinite horizontal planes with spacing H. The hydraulic diameter equivalence ( ), i.e., the height depicted in Supplementary Fig. 6 corresponds to one half of the diameter of circular cross-section, i.e., H = 0.5D. Therefore, Eq. (4) for a cylindrical enclosure featuring a thermo-electrochemical cell of a cylindrical enclosure is given by: The variables and thermal properties needed to calculate the Rayleigh number for the 0.4M potassium ferri/ferrocyanide as electrolyte are given in Supplementary Table 2.
Here, the density and kinematic viscosity values of the electrolyte are estimated using the temperature-dependent formula given in the literature 3,4 .
Utilizing the values in the table, the reciprocal of quadratic root of Rayleigh number ( ) is calculated to be 0.11 and the ratio of height to length ( ) is 0.06. These values satisfy Eq. (5) or , which is the criterion for negligibly small heat transfer to occur because of the aforementioned reasons. This result implies that the convection heat transfer from the heated electrode directing to the cooled electrode is