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A study on the effect of oxidation on critical heat flux in flow boiling with downward-faced carbon steel
International Journal of Heat and Mass Transfer 147 (2020) 118966
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A study on the effect of oxidation on critical heat flux in flow boiling with downward-faced carbon steel Kai Wang a,⇑, Nejdet Erkan a, Koji Okamoto b a b
Department of Nuclear Engineering and Management, School of Engineering, The University of Tokyo, 7-3-1 Hongo, Bunkyo-ku, Tokyo 113-8654, Japan Nuclear Professional School, School of Engineering, The University of Tokyo, 2-22 Shirakata, Tokai-mura, Ibaraki 319-1188, Japan
a r t i c l e
i n f o
Article history: Received 15 March 2019 Received in revised form 16 August 2019 Accepted 26 October 2019
Keywords: Critical heat flux Flow boiling Bubble behavior Heat transfer Oxidation effect
a b s t r a c t To execute in-vessel retention external reactor vessel cooling (IVR-EVRC) successfully, critical heat flux (CHF) plays a key role in securing the thermal and structural integrity of the reactor pressure vessel (RPV). In real-world applications, the RPV outer surface is exposed and oxidized. In this study, we performed a downward-face flow boiling CHF experiment using a carbon steel plate with several boiling cycles under water. Because the direction of gravity was inverse with the normal of the wall, it was difficult to remove bubbles, and in the local area, some local burning could repeat several times. After polishing the surface with sandpaper, the heat was supplied to the surface by cartridge heaters in a step-by-step manner until CHF was reached, after which the surface was cooled down. Using the same surface, without any further surface treatment, this heating cycle was repeated five times. The CHF value increased with the number of cycles. When the experiment was repeated for different mass fluxes, the CHF value also increased with the number of cycles. We suggest that this increasing tendency in CHF is associated with the increasing level of surface oxidation. After every repetition of the heating cycle, surface oxidation level increased, fewer bubbles were observed on the surface, resulting in higher CHF values. After the experiment, surface roughness increased and contact angle decreased. The liquidvapor mixture area decreased with increasing heat flux in one experiment, whereas it decreased with boiling time, under the same heat flux, which delayed CHF. The gradual oxidation process of carbon steel could be beneficial for real-world applications of IVR-ERVC. Ó 2019 Elsevier Ltd. All rights reserved.
1. Introduction In-vessel retention external reactor vessel cooling (IVR-EVRC) is a severe accident management strategy in which vessel breaching is prevented by flooding the reactor pressure vessel (RPV) cavity with water, so decay heat generated by the fuel rods can be removed by heat conduction through the RPV wall [1]. In this situation, critical heat flux (CHF) is likely to cause RPV failure by weakening the mechanical strength [2]. Consequently, CHF has been a common research focus in both pool and flow boiling experiments. Copper is widely used for pool boiling experiments, owing to its high thermal conductivity [3,4]. However, different materials have different CHF [5], and in real-world situations, lower plenum RPVs are composed of carbon steel. By performing pool boiling experiments with SA508, Kam et al. [6] found that as the width size decreased, the CHF enhancement ratio increased for carbon steel ⇑ Corresponding author at: 7-3-1 Hongo, Bunkyo-ku, Tokyo 113-8654, Japan. E-mail address: [email protected] (K. Wang). https://doi.org/10.1016/j.ijheatmasstransfer.2019.118966 0017-9310/Ó 2019 Elsevier Ltd. All rights reserved.
compared with stainless steel, especially at high inclination angles; they concluded that such factors as static contact angles and dissolved oxygen concentrations were not sufficient to explain the difference. Lee and Chang [7] conducted pool boiling CHF experiments with SA508 and SS304 test heaters under atmospheric pressure; they found higher CHF for SA508 and concluded that magnetite (Fe3O4) on the heater surface was produced by corrosion. Some studies have also been focused on flow boiling. Chang et al. [8] conducted flow boiling CHF experiments using a full scale 2D curved test section under atmospheric pressure with a SA508 Gr3 carbon steel heater; the results showed that SA508 displayed different (higher) CHF behavior in comparison with stainless steel, copper, and aluminum. Azzian et al. [9] tested carbon steel in a flow boiling loop designed specifically for IVR conditions. Up to 70% CHF enhancement values was observed for oxidized carbon steel in comparison with stainless steel. Park et al. [10] considered the effect of the heater material on CHF, finding higher CHF values for SA508 in comparison with stainless steel in all cases; they concluded that this behavior reflected the oxidation of carbon steel in the aqueous environment. Trojer et al. [11] showed that the
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pre-oxidized carbon steel had 70% CHF enhancement compared to that of the stainless steel in the flow boiling. Regarding other materials, Lee et al. [12] performed a saturated pool boiling experiment using zircaloy oxidized at 300, 450, and 600 °C, and increases in the CHF and wettability were found after the surface was oxidized. Lee et al. tried to explain the 40% CHF enhancement by the decreasing contact angle. Vlachou et al. [13] found that copper could also be oxidized and that this type of oxidation could decrease the heat transfer coefficient in the flow boiling. Coursey and Kim [14] obtained enhancements in both the CHF and wettability of copper after oxidation. Tachibana et al. [15] reported a CHF enhancement of oxidized aluminum CHF compared to that of the stainless steel. While most researchers have observed an increase in CHF with oxidation, Son et al. [16] found an inverse effect in which the CHF of bare SA508 was approximately 140% of Zuber’s pool boiling CHF prediction [17], whereas 3, 10, 20, 30, and 40 days of oxidation reduced the CHF of SA508 to approximately 65% of the Zuber’s prediction in the pool boiling.
Given the important role played by carbon steel in IVR, we investigated the CHF of carbon steel during downward face flow boiling instead of upward face boiling. The reason for conducting downward face boiling is that the lower head of RPV has a downward surface and the direction of gravity is inverse to the wall normal, so it is difficult for the bubbles generated on the lowest part to escape and CHF may occur earlier. In Wang et al. [18], a pool boiling experiment using carbon steel was carried out to investigate the CHF change due to local boiling. In the downward face boiling, it is very difficult for the bubbles to escape, so they may linger for a longer time. Thus, it is very likely that local burning may occur in some regions. This type of local boiling probably resulted from non-uniform heat flux around the lower head. The natural convection in the lower head pool caused by the temperature gradient between the pool boundary layer and the middle of the pool resulted in the non-uniform heat flux distribution.[19] In addition, the emission of fission gas and falling down of the melting fuels could cause the temperature and heat flux distribution to be unsteady. If at some point, the heat flux is greater
(a) Schematic of the experimental apparatus
(b) A photograph of the experimental apparatus Fig. 1. The experimental apparatus.
K. Wang et al. / International Journal of Heat and Mass Transfer 147 (2020) 118966
than the upper limit, CHF may happen. The heat flux may vary with time, making the heat flux of the local area change constantly. Thus, some points can experience multiple CHF situations. In Wang et al. [18] it was found that the CHF increased after several cycles in the pool boiling situations. However, few researchers have focused on the effect of flow boiling local burning on CHF. This type of local boiling may occur several times, so the local surface morphology can change through different cycles of local boiling. The boiling process and CHF will also change owing to the surface morphology change. Thus in this study, the local burnout of the downward face flow boiling was investigated.
2. Experimental setup and procedure 2.1. Experimental apparatus The schematic of the flow boiling experimental apparatus is shown in Fig. 1. The main components were the test section, three thermocouples in intervals of 3 mm for measuring the temperature of the copper block base, two high-speed cameras, one upstream tank with a thermocouple to measure the upstream temperature, and one downstream tank with a thermocouple to measure the downstream temperature. A pump with a maximum volume flow rate of 3 m3/h of distilled water was installed near the downstream tank. One electromagnetic flowmeter was applied to measure the
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volume flow rate. Distilled water was supplied to the flow channel from the upstream tank. The flow channel was composed of acrylic transparent glass for visualization and had a cross section of 40 mm (width) by 10 mm (height). The test section was installed from the top of the channel to reach a downward heating surface. A reflux condenser was installed at the top of the downstream tank to condense overheated steam. Two identical high-speed cameras (Photron Fastcam SA5) were used. These cameras can operate with a maximum resolution of 1024 1024, 7000 fp. One camera was placed next to the water tank to capture images of the downward side reflected from a mirror under the water tank, and the other was placed near the water tank to obtain a side view. The two cameras were synchronized by connecting them together using the input/output function. A data logger connected to the five thermocouples was used for temperature measurement. The test section was composed of a copper base, carbon steel plate (heating zone), and polyether ether ketone (PEEK; thermal isolation zone) as shown in Fig. 2. The copper base was 40 40 mm rectangular on top and 30 30 mm on the bottom. The boiling surface of the carbon steel plate used here had an area of 30 30 mm rectangular. In total, nine cartridge heaters (each with a 225 W power rating) were inserted at the top of the copper block base controlled by a slidac. The reason that soldering method was adopted instead of using a whole carbon steel block was that the shielding material PEEK has a very low melting temperature, and the silicone used between
(a) Test section schematic
(b) A photograph of the test section Fig. 2. Test section schematic.
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the copper and PEEK cannot sustain temperatures greater than 200 °C. Moreover, the thermal conductivity of carbon steel is relatively low. High heat flux cannot be obtained by direct heating of the carbon steel block because the temperature would exceed the PEEK melting temperature; therefore, the carbon steel plate was attached to a copper using the silver soldering method. A similar method was also used by Mei et al. [20], who employed silver solder to adhere both carbon steel and stainless steel to the copper base. Detailed geometrical information on the copper block base is shown in Fig. 2. The carbon steel plate was approximately 1.0 mm. Golobicˇ and Bergles [21] found that when the heater surface exceeded a certain thickness, the CHF was stable. The asymptotic thickness is related to the thermal effusivity and for the carbon steel, and it was calculated to be approximately 0.36 mm, which is less than the thickness of the carbon steel plate (about 1.0 mm) used in this study. Thus, the effect of thickness could be eliminated in this study. After soldering, we measured the thermal conductivity of the silver solder layer. A thermocouple was attached to the bottom of the carbon steel surface and heat flux was applied by a slidac. Then, the temperature profile was recorded. The solder layer was measured to be approximately 0.1 mm and had a thermal conductivity of 15.0 W=ðm KÞ. The heat transfer through the heater block could be simplified using 1D steady-state heat conduction. The temperature of a carbon steel surface can be calculated by using Eq. (1), where T1 is the temperature of the copper base wall calculated using the 1D Fourier law,dcu ¼ 2 mm,kcu ¼ 400 W=ðm KÞ,dcs ¼ 1:0 mm, kcs ¼ 50 W=ðm KÞ [22], dsolder ¼ 0:1mm; ksolder ¼ 15:0 W=ðm KÞ:
T wall ¼ T 1 q
00
dcs dcu dsolder þ þ kcs kcu ksolder
ð1Þ
where dcu , dcs , and dsolder are the length of the copper, carbon steel, and silver solder layer, respectively, and kcu , kcs , and ksolder are the thermal conductivity of the copper, carbon steel, and solder layer, respectively. After oxidation, a layer of oxidized particles will be generated on the heating surface. The estimated layer thickness was 0.3 lm for a surface oxidized for 3 days [16], which is quite small compared to the thickness of the carbon steel plate. Thus, the oxidation effect on the temperature change of the wall was 00 not considered.q is the heat flux rate, which can be calculated using Eq. (2), and x denotes the distance between the copper thermocouples. In Jeong and Kim [23], the temperature gradient was also calculated based on the three thermocouples. They adopted a three-point, backward-space Taylor series approximation. In this study, the temperature gradient was calculated using a threepoint least-square fittings, as shown in Eq. (3). 00
q ¼ kcu
dT dx
ð2Þ
P P P dT ð3 xi T i xi T i Þ ¼ P 2 P dx 3 xi ð xi Þ2
ð3Þ
Here, xi denotes the distance from the copper base down face, and i = 1, 2, 3.T i is the temperature of the corresponding thermocouple. The total heat loss during the experiment was assessed using Eq. (4)
ahl ¼
V2 R
00
q S
ð4Þ
2
V R
00
where S is the area of the boiling surface, q is the heat flux derived from the thermocouple temperature difference, V is the voltage value measured by the multimeter connected to the slidac, and R is the resistance of the cartridge heater. The average heat loss percentage ahl was estimated to be approximately 4.9%, which includes both the heat dissipated to the air through the copper base or cartridge heater and the heat dissipated to the water through the silicon insulation layer.
2.2. Experimental procedure Owing to the easiness of carbon steel oxidation, the surface was completely cleaned prior to the experiment. Three types of sandpapers were used to clean the surface. To wipe off all residuals on the surface, very coarse P80 sandpaper was used firstly; then, fine P600 and P1200 sandpapers were used to polish the surface; and finally, the surface was cleaned by acetone. Before the experiment, we heated up the distilled water to its saturated temperature and the temperature was kept by controlling the heater. The air was evacuated for 30 min. It took approximately 2 h for heating and air evacuation. Therefore, before each experiment, the test piece had already been in the circulating water for approximately 2 h. During the experiment, the power of the cartridge heaters was increased gradually until CHF was reached. The voltage was increased by 20 V firstly, then 10 V, and 5 V when it came close to CHF, which was about 0.03 to 0.06 Mw/m2 of heat flux. The temperature increased and became stable at a certain point, and if the temperature was maintained for at least 1 min, we recorded this point and increased the heat flux for another point. CHF conditions were assumed when the temperatures of the thermocouples suddenly increased significantly (i.e., more than 50 °C over a very short time), and also at this time, a very long vapor film will form along the heating surface. The power supply was then stopped to prevent further damage. After CHF data were obtained and the power supply stopped, the temperature of both the water and block decreased. When the temperature of the block dropped to less than 95 °C, power was re-supplied until the next CHF value was
Table 1 Experimental uncertainties. Parameter
Measuring unit
Range
Uncertainty
Fluid temperature Inlet temperature Wall temperature Length Mass flux (160 kg/(m2 s)) Mass flux (320 kg/(m2 s)) Mass flux (640 kg/(m2 s)) CHF at 160 kg/(m2s) CHF at 320 kg/(m2s) CHF at 640 kg/(m2s)
K-type thermocouple N/A K-type thermocouple N/A Electromagnetic flowmeter Electromagnetic flowmeter Electromagnetic flowmeter
243.5–700 K N/A 243.5–700 K 0–3 mm 0–40 L/min 0–40 L/min 0–40 L/min
±0.75% 1% ±0.75% 3.3% 5% 2.5% 1.25% 6.1% 4.3% 3.7%
K. Wang et al. / International Journal of Heat and Mass Transfer 147 (2020) 118966
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reached. Cycles were repeated five times for the same surface, without any surface cleaning between cycles. 2.3. Uncertainty analysis The estimated uncertainty of the heat flux was obtained using Eq. (5) [24]:
dR ¼ R
sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi du1 2 du2 2 dun 2 ð Þ þ ð Þ þ þ ð Þ u1 u2 un
ð5Þ
Table 2 Experimental flow rate parameters. Mass flux (kg/(m2 s)
Volume flow rate (L/min)
Velocity (m/s)
Reynolds number
160 320 640
4 8 16
0.17 0.33 0.67
0.91 104 1.82 104 3.64 104
(a) 160 kg/(m2 s)-A
Fig. 4. CHF enhancement with the boiling cycle for different mass fluxes.
(b) 160 kg/(m 2 s)-B
(c) 320 kg/(m2 s)-A
(d) 320 kg/(m2 s)-B
(e) 640 kg/(m2 s)-A
(f) 640 kg/(m2 s)-B
Fig. 3. Boiling curves for carbon steel plate. Experiment under a mass flux of (a) 160 kg/(m2 s)-A, (b) 160 kg/((m2 s))-B, (c) 320 kg/((m2 s))-A, (d) 320 kg/((m2 s))-B, (e) 640 kg/((m2 s))-A and (f) 640 kg/((m2 s))-B.
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Fig. 5. Pre- and post-experimental images of the carbon steel surface. (a) Pre-experimental surface. (b) Submerged in water. Experiment using a mass flux of (c) 160 kg/(m2 s), (d) 320 kg/((m2 s)), and (e) 640 kg/((m2 s)).
In this approach, the dependent variable R has independent linear parameters, such as R = R (u1, u2. . .un). The experimental uncertainties are reported in Table 1. The volume flow rate of the electromagnetic flowmeter is 0–40 L/min, and has ±0.2 L/min error range. This will result in a bigger uncertainty under smaller mass flux. 3. Results and discussion To investigate the enhancement effect of oxidation under different mass fluxes, three different mass fluxes were applied: 160, 320, and 640 kg/(m2 s). The relationship between volume flow rate V_ and mass flux G is
V_ ¼ GLH=ql Fig. 6. Average CHF as a function of mass flux. An empirical CHF correlation (Katto and Kurata [27]) is provided for comparison.
ð6Þ
where L (=40 mm) is the channel width, and H (=40 mm) is the channel height.ql (=958 kg/m3) is the density of water at saturation temperature.
Fig. 7. Images of the carbon steel surface during experiments under different mass fluxes and the same heat flux (=0.09 MW/m2). Experiments conducted at (a) 160 kg/(m2 s)1, (b) 160 kg/(m2 s)-2, (c) 160 kg/(m2 s)-3, (d) 160 kg/(m2 s)-4, (e) 160 kg/(m2 s)-5. Experiments conducted at (f) 320 kg/(m2 s)-1, (g) 320 kg/(m2 s)-2, (h) 320 kg/(m2 s)-3, (i) 320 kg/(m2 s)-4, and (j) 320 kg/(m2 s)-5. Experiments conducted at (k) 640 kg/(m2 s)-1, (l) 640 kg/(m2 s)-2, (m) 640 kg/(m2 s)-3, (n) 640 kg/(m2 s)-4, and (o) 640 kg/(m2 s)-5.
K. Wang et al. / International Journal of Heat and Mass Transfer 147 (2020) 118966
The relationship between velocity
v ¼ G=ql
v
and mass flux G is
ð7Þ
Reynolds number
Re ¼ GD=l
ð8Þ
Where l (=0.28 mPa s) is the viscosity of the water at saturation temperature, D (=16 mm) is the hydraulic diameter of the channel. Their velocities and Reynolds numbers are listed in Table 2. 3.1. Experimental boiling curves The boiling curves obtained in the experiments are shown in Fig. 3. For each experiment, the results from five CHF tests were obtained, and recorded as m kg/(m2 s)-n, where m is the mass flux in kg/(m2 s) and n is the cycle number. Experiments under different mass flux were conducted twice and was denoted as A and B. Generally, under low heat flux, the superheat increases with boiling cycle, indicating a deterioration of heat transfer. When the heat flux is close to CHF, the endpoint increases, confirming that the superheating and CHF increase with boiling time. The first CHFs of each mass flux were almost the same, indicating a good initial surface morphology control. However, after the first cycle,
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the CHF enhancement became rather random due to the randomness of the oxidation process. In Wang et al. [18], experiments were done for 3 times at the same condition, in which the initial CHFs were the same in each experiment, but the CHF enhancement showed different ratios due to the randomness of the oxidation process. Although the experiments were done for a long time, it might not reach a saturate CHF. The reasons are as follows: After 5 cycles, the CHF has not reached a stable point. The oxidation process is a random process, and CHF enhancement is also different in each experiment. Fig. 4 shows the CHF enhancement of each experiment. In all the experiments, CHF showed different enhancement with boiling cycle. Images captured before and after each experiment show the condition of the surface (Fig. 5). The initial polished surface contained only a few scratches (Fig. 6a). After the experiment, the surface was immediately taken out from the water. Some surface oxidation occurred after all the experiments (Fig. 6c, 6d, and 6e). The surface roughness has been measured using a portable surface roughness tester SJ-210. After the test section was submerged in water for about 2 h, the test section showed a very slight increase in surface roughness as shown in Fig. 6b, which suggested that the surface received small amount of oxidation. After the experiment,
Fig. 8. Bubble film departure with a mass flux of 320 kg/(m2 s)-5 for different heat flux levels. Heat flux = 0.33 MW/m2: (a) down view, (b) side view; Heat flux = 0.81 MW/m2: (c) down view, (d) side view; post CHF (=0.87 MW/m2): (e) down view and (f) side view.
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the surface roughness was also measured and all the test surfaces had an increase in surface roughness. The increase of surface roughness was found in Son et al. [16] after oxidizing carbon steel in air for a long time. However, it was reported that the surface roughness had little effect on CHF. Hahne and Grigull [25] pointed out that the surface roughness might influence the vapor bubble growth but it became weaker near the peak heat flux. O’Hanley et al. [26] examined the separate effects of surface wettability, porosity, and roughness on the CHF using engineered surfaces, and concluded that surface roughness had no effect on CHF within the limit of the database. Several empirical and semi-empirical models have been proposed for CHF determination. Unfortunately, most of the correlations are based on vertical tube experiments. We chose an empirical CHF correlation (Katto and Kurata [27]) for comparison, which can be expressed as
qCHF ¼ 0:186Ghfg
qg qf
!0:559
rqf 0:264 G2 L
ð9Þ
whereG is the mass flux of the liquid, and L is the length of the heated surface. The correlation was developed based on vertical flow boiling experiments with a heating surface of 10 to 20 mm in length for saturated water and R-113. Comparison results are shown in Fig. 6. The first cycle CHF values all fell under the Katto and Kurata’s equation line; this is because the Katto and Kurata’s equation was derived from upward
flow boiling whereas our experiments were downward-face flow boiling. Downward-face flow boiling has a relatively smaller CHF owing to the difficulty of vapor removal because it was not easy for the bubbles generated on the downward face to escape from the surface due to gravity. An attempt was made to predict the CHF of downward face based on the Katto and Kurata’s equation. The modification was made mainly to the exponent of the term
rqf G2 L
, which is the ratio
of inertial force to surface tension force. The modification was done using a least-squares fitting method, and the results (Eq. (10)) showed good agreement with the results.
qCHF ¼ 0:235Ghfg
qg qf
!0:559
rqf 0:391 G2 L
ð10Þ
4. Discussion Detailed images of the downward surface were collected during the experiments (Fig. 7). It is observed that although not many, still a few bubbles came out from the heater sides. The bubbles generated on the edges showed some heat might be dissipated through the silicon layer between the PEEK and carbon steel plate, which was part of the heat loss explained in Eq. (4) in chapter 2. With the same heat flux, a longer oxidation time resulted in fewer but larger bubbles on the surface. The surface also became
Fig. 9. Surface images under different heat fluxes for an experiment with a mass flux of 160 kg/(m2 s)-5. Heat flux values of (a) 0.18 MW/m2, (b) 0.31 MW/m2, (c) 0.47 MW/m2, and (d) 0.68 MW/m2.
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progressively darker, indicating increased level of oxidation of the carbon steel during the boiling process. Greater oxidation was consistent with fewer but larger bubbles. Studies on the impact of the nucleation site density (NSD) on the heat transfer coefficient and CHF have yielded various results. For increasing CHF, Jaikumar et al. [28] fabricated the microstructure and performed a pool boiling experiment, which showed that both the heat transfer coefficient and CHF increased due to the change of NSD and wettability. Kurihara and Myers [29] reported that the nucleate boiling enhancement was due to an increase in active NSD. Betz et al. [30] found that hydrophobic surfaces can delay CHF. Chang and You [31] showed that that increasing active nucleation sites increased pool boiling heat transfer coefficients by 30%, and
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increases CHF by approximately 100%. On the contrary, You et al. [32] found that nanoparticles in water decrease the number of bubbles and increased the CHF. Yeom et al. [33] reported a reduction in the NSD and an increase in the CHF in the pool boiling experiment using nanoparticle coatings formed by electrophoretic deposition of Ti and TiO2 particles. Wang et al. [34] found an inverse relationship between NSD and CHF by irradiating a copper block. They concluded that greater NSD generates more bubbles and increases the heat transfer coefficient; furthermore, the area of bubble vapor covering the surface increases, which prevents water supply to the boiling surface and results in the early occurrence of CHF. Wang et al. [18] also reported that the CHF was enhanced with oxidation and fewer bubbles generated on the sur-
Fig. 10. Surface images during subsequent heating cycles under similar heat flux and a mass flux of 160 kg/(m2 s). Image of (a) cycle 1, 0.37 MW/m2, (b) cycle 2, 0.33 MW/m2, (c) cycle 3, 0.32 MW/m2, (d) cycle 4, 0.32 MW/m2, and (e) cycle 5, 0.31 MW/m2.
Fig. 11. Contact angle of the carbon steel surface. (a) Before experiment, (b) Submerged in water. After the experiment under a mass flux of (c) 160 kg/(m2 s), (d) 320 kg/(m2 s), (e) 640 kg/(m2 s).
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face in the pool boiling. Wang et al. [35] observed both an increase and a decrease of CHF due to the change in NSD. They identified a threshold NSD (<~5 105 sites/m2), above which the CHF was deteriorated due to the increase in the NSD and vice versa. Our experimental results are consistent with those of a pool boiling experiment conducted by Wang et al. [18], in that a decrease in NSD delayed the CHF. This delay probably occurred because in the downward face boiling, it is easy for bubbles to accumulate due to the difficulty of bubble removal, so more bubbles will delay CHF occurrence. We found that under relatively small heat flux, bubbles larger than a certain size moved and left the surface, but under medium–high heat flux, small bubbles coalesced to form a bubble film. Using images captured by high-speed cameras, we could easily distinguish that under medium heat flux, this bubble film was able to move off the surface. Fig. 8 shows bubble film departure for different heat fluxes. According to the side view camera, most of the bubble film generated on the surface did not exceed the channel height, which meant that channel height would not restrict the bubble behavior. Actually, Katto and Kurata [27] used a rectangular nozzle with a width of 15 mm, and two heights 10 mm mainly and 5 mm auxiliary. And they found that difference of CHF was hardly perceived for the change of channel height from 10 mm to 5 mm. For a medium heat flux (Fig. 8a and b), the bubble film departs the surface without connecting to the next bubble film, which means that little vapor was produced. For high heat flux (Fig. 8c and d), the departing bubble film has some connections to neighboring bubble films, which means some water was prevented from supplying to the surface. Fig. 8e and f show CHF images; the bubble film extends continuously across the surface, showing that as the heat flux increases, bubble vapor continues and no water can be supplied for heat removal. Fig. 9 shows surface images under different heat fluxes with a mass flux of 160 kg/(m2 s)-5. Initially, several small bubbles coexisted with liquid (i.e., the liquid–vapor mixture area); however, as heat flux increased, this area decreased and more vapor was generated to cover the surface more, preventing water from reaching to the surface and thus leading to CHF. Fig. 10 shows surface images of each experimental cycle with a mass flux of 160 kg/(m2 s) and a constant heat flux. There is an increase in the mixture area with time, resulting in the delay of CHF appearance, which probably reflects a decrease in the NSD. This phenomenon was also observed by Gong et al. [36], who used a copper surface and found that after irradiation, the liquid-vapor mixture area was enlarged compared with the non-irradiated surface. They attributed this enlargement to a change in hydrophilicity after irradiation. A decrease in the contact angle (i.e., an increase in the wettability) after oxidation has been observed in various studies [6,12,16], and was also found in our research (Fig. 11). The contact angle was measured before and after the experiment. It showed clearly that the contact angle decreased after the experiment. Mori and Utaka [37] noted that CHF enhancement may be due to a change of NSD, wettability and capillary wicking. Lee et al. [12] explained the 40% CHF enhancement of oxidized zircaloy by the increasing wettability. However, the contact angle alone cannot explain the CHF enhancement, as recently reported by Son et al. [38], who stated that nanostructure, nanoporous, and microstructure surfaces exhibited CHF enhancement compared to a bare surface, as well as smaller contact angles and increased liquid wicking. In this experiment, it is difficult to quantify the wicking effect due to the non-uniform surface change. Nevertheless, the change in contact angle and surface morphology indicated that the surface property change did influence flow boiling, but it was difficult to identify whether it was the wicking effect.
To confirm these surface changes, energy dispersive spectroscopy (EDS) of the oxidized surface was performed. Due to the difficulty of conducting EDS experiments on the surface after experiment, small round test pieces heating in the water under saturation temperature to simulate the actual experimental environment were observed instead. Fig. 12 showed the EDS results of the surface which were non-oxidized, 1 h oxidized in water and 3 h oxidized in water, respectively. For all the test specimens, the peak of iron, oxygen and carbon were detected. The mass content of oxygen increased from 1.09%, 2.90% to 5.90%, which clearly showed iron oxidation. In the non-oxidized specimen, very small amount of oxygen was detected, which might be some oxygen in air environment. The oxygen content was also detected in Chi et al. [12], which the non-oxidized zircaloy surface also contained small amount of oxygen. The non-oxidized and oxidized carbon steel surfaces were also imaged by backscattered electron imaging using an SEM (Fig. 13). Scratches and small cavities are evident on the surface without oxidation in Fig. 13a and 13b, and some oxidation particles had
(a) Original test piece
(b) 1 h oxidized test piece
(c) 3 h oxidized test piece Fig. 12. Surface EDS of (a) original test piece, (b) 1 h oxidized test piece and (c) 3 h oxidized test piece.
K. Wang et al. / International Journal of Heat and Mass Transfer 147 (2020) 118966
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Fig. 13. SEM images of the non-oxidized and oxidation carbon steel. Images acquired at (a) 4000 magnification for non-oxidized carbon steel, (b) 10000 magnification for non-oxidized carbon steel, (c) 1000 magnification for oxidized carbon steel, and (d) 8000 magnification for oxidized carbon steel.
Fig. 14. Schematic of the gradual oxidation of a carbon steel surface.
occurred within these scratches, covering the cavities and decreasing the NSD in Fig. 13c and 13d. This mechanism for the gradual oxidation of the carbon steel surface is illustrated in Fig. 14. In fact, the filling of the pores and cavities has also been reported. For example, Paz et al. [39] carried out subcooled flow boiling experiments using copper test section with different roughness, and noted that the surface morphology should be a very important factor but has usually been ignored in correlations and experiments. Lee et al. [40] also noted that the decrease in the heat transfer coefficient was due to the surface morphology change
caused by the formation of aluminum hydroxide on the boiling surfaces. This process causes the pores between the nanostructures to be filled and smoothed.
5. Conclusions In this study, flow boiling CHF experiments were conducted on a downward facing surface under different mass fluxes to investigate the effect of oxidation. The main conclusions are as follows:
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K. Wang et al. / International Journal of Heat and Mass Transfer 147 (2020) 118966
(1) During the boiling process, the CHF increased with boiling cycle gradually as fewer bubbles generated on the heating surface. (2) After the experiment, surface roughness became greater than the original test surface. (3) The correlation based on the Katto and Kurata’s equation for downward-face agreed well with the experimental results. (4) The liquid-vapor mixture area decreased with increasing heat flux in the experiment. Furthermore, with increasing boiling time, the mixture area decreased under a constant heat flux, which delayed the CHF. (5) The increasing CHF contributed to increasing the wettability and decreasing NSD. In real-world applications, RPV surfaces tend to corrode with time, and the changing surface characteristics impact the boiling phenomena and CHF significantly. In the real RPV lower head, because the direction of gravity is inverse with the normal of the wall, bubbles are difficult to be removed and in some area, some local burning may repeat several times. Our results demonstrate the advantages of using EVRC to dissipate the residual heat that the CHF increases with the local burnout. However, additional research is necessary to examine whether this type of enhancement will continue to increase or saturate with further cycle repetition. Declaration of Competing Interest We declare that we have no conflict of interest. Acknowledgements This work is mainly supported by the Japan Grant-in-Aid for Scientific Research (B) (Grand No. 19H02645). The author would like to thank Dr. Huilong Yang from Nuclear Professional School, School of Engineering, The University of Tokyo for his help on the SEM and EDS analyses. References [1] J.L. Rempe, K.Y. Suh, F.B. Cheung, S.B. Kim, In-vessel retention of molten corium: lessons learned and outstanding issues, Nucl. Technol. 161 (3) (2008) 210–267. [2] T.G. Theofanous, C. Liu, S. Additon, S. Angelini, O. Kymaelaeinen, T. Salmassi, In-vessel coolability and retention of a core melt, Nucl. Eng. Des. 169 (1–3) (1997) 1–48. [3] M.S. El-Genk, Z. Guo, Transient boiling from inclined and downward-facing surfaces in a saturated pool, Int. J. Refrig. 16 (6) (1993) 414–422. [4] M.S. El-Genk, A.F. Ali, Enhanced nucleate boiling on copper micro-porous surfaces, Int. J. Multiph. Flow. 36 (10) (2010) 780–792. [5] K. Wang, N. Erkan, H. Gong, L. Wang, K. Okamoto, Comparison of pool boiling CHF of a polished copper block and carbon steel block on a declined slope, J. Nucl. Sci. Technol. 9 (55) (2018) 1–14. [6] D.H. Kam, Y.J. Choi, Y.H. Jeong, Critical heat flux on downward-facing carbon steel flat plates under atmospheric condition, Exp. Therm. Fluid Sci. 90 (2018) 22–27. [7] J. Lee, S.H. Chang, An experimental study on CHF in pool boiling system with SA508 test heater under atmospheric pressure, Nucl. Eng. Des. 250 (2012) 720–724. [8] H. Chang, T. Hu, W. Lu, S. Yang, X. Zhang, Experimental study on CHF using a full scale 2-D curved test section with additives and SA508 heater for IVR-ERVC strategy, Exp. Therm. Fluid Sci. 84 (2017) 1–9. [9] R. Azizian, T. McKrell, K. Atkhen, J. Buongiorno, Effects of porous superhydrophilic surfaces on flow boiling critical heat flux in IVR accident scenarios, American Nuclear Society, Chicago, 2015. [10] H.M. Park, Y.H. Jeong, S. Heo, Effect of heater material and coolant additives on CHF for a downward facing curved surface, Nucl. Eng. Des. 278 (2014) 344– 351.
[11] M. Trojer, R. Azizian, J. Paras, T. McKrell, K. Atkhen, M. Bucci, J. Buongiorno, A margin missed: The effect of surface oxidation on CHF enhancement in IVR accident scenarios, Nucl. Eng. Des. 335 (2018) 140–150. [12] C.Y. Lee, T.H. Chun, W.K. In, Critical heat flux of oxidized zircaloy surface in saturated water pool boiling, J. Nucl. Sci. Technol. 52 (4) (2015) 596–606. [13] M.C. Vlachou, J.S. Lioumbas, K. David, D. Chasapis, T.D. Karapantsios, Effect of channel height and mass flux on highly subcooled horizontal flow boiling, Exp. Therm. Fluid Sci. 83 (2017) 157–168. [14] J.S. Coursey, J. Kim, Nanofluid boiling: the effect of surface wettability, Int. J. Heat Mass Transf. 29 (2008) 1577–1585. [15] F. Tachibana, M. Akiyama, H. Kawamura, Non-hydrodynamic aspects of pool boiling burnout, J. Nucl. Sci. Technol. 4 (3) (1967) 121–130. [16] H.H. Son, U. Jeong, G.H. Seo, S.J. Kim, Oxidation effect on the pool boiling critical heat flux of the carbon steel substrates, Int. J. Heat Mass Transf. 93 (2016) 1008–1019. [17] N. Zuber, Hydrodynamic aspects of boiling heat transfer (thesis), California. Univ., Los Angeles; and Ramo-Wooldridge Corp., Los Angeles, 1959. [18] K. Wang, N. Erkan, H. Gong, K. Okamoto, Effects of carbon steel surface oxidation on critical heat flux in downward-face pool boiling, Int. J. Heat Mass Transf. 136 (2019) 470–485. [19] B.R. Sehgal, Nuclear safety in light water reactors: severe accident phenomenology, second ed., Elsevier, San Diego, 2012. [20] Y. Mei, Y. Shao, S. Gong, Y. Zhu, H. Gu, Effects of surface orientation and heater material on heat transfer coefficient and critical heat flux of nucleate boiling, Int. J. Heat Mass Transf. 121 (2018) 632–640. [21] I. Golobicˇ, A.E. Bergles, Effects of heater-side factors on the saturated pool boiling critical heat flux, Exp. Therm Fluid Sci. 15 (1) (1997) 43–51. [22] Engineering Toolbox, Thermal conductivity of common materials and gases, 2003. https://www.engineeringtoolbox.com / thermal-conductivity-d_429. html. [23] U. Jeong, S.J. Kim, Subcooling effect on boiling heat transfer of inclined downward-facing surface under low flow and pressure, Int. J. Heat Mass Transf. 127 (2018) 182–195. [24] R.J. Moffat, Describing the uncertainties in experimental results, Exp. Therm. Fluid Sci. 1 (1) (1988) 3–17. [25] E. Hahne, U. Grigull. Heat transfer in boiling. Hemisphere, New York (1977). [26] H. O’Hanley, C. Coyle, J. Bungiorno, T. McKrell, L.W. Hu, M. Rubner, R. Cohen, Separate effects of surface roughness, wettability and porosity on the boiling critical heat flux, Appl. Phys. Lett. 103 (2013) 024102. [27] Y. Katto, C. Kurata, Critical heat flux of saturated convective boiling on uniformly heated plates in a parallel flow, Int. J. Multiphase Flow 6 (6) (1980) 575–582. [28] A. Jaikumar, A. Rishi, A. Gupta, S.G. Kandlikar, Microscale morphology effects of copper–graphene oxide coatings on pool boiling characteristics, J. Heat Transf. 139 (2017) 111509. [29] H.M. Kurihara, J.E. Myers, The effects of superheat and surface roughness on boiling coefficients, AIChE J. 6 (1) (1960) 83–91. [30] A.R. Betz, J. Jenkins, D. Attinger, Boiling heat transfer on superhydrophilic, superhydrophobic, and superbiphilic surfaces, Int. J. Heat Mass Transf. 57 (2) (2013) 733–741. [31] J.Y. Chang, S.M. You, Enhanced boiling heat transfer from microporous surfaces: effects of a coating composition and method, Int. J. Heat Mass Transf. 40 (18) (1997) 4449–4460. [32] S.M. You, J.H. Kim, K.H. Kim, Effect of nanoparticles on critical heat flux of water in pool boiling heat transfer, Appl. Phys. Lett. 83 (16) (2003) 3374–3376. [33] H. Yeom, K. Sridharan, L.M. Corradini, Bubble dynamics in pool boiling on nanoparticle-coated surfaces, Heat Transfer Eng. 36 (2015) 1013–1027. [34] L. Wang, N. Erkan, H. Gong, K. Okamoto, Electron beam irradiation effect on critical heat flux in downward-facing flow boiling, Int. J. Heat Mass Transf. 120 (2018) 300–304. [35] K. Wang, S. Gong, B. Bai, W. Ma, On the Relation between nucleation site density and critical heat flux of pool boiling, Heat Transf. Eng. (2017) 1–9. [36] H. Gong, L. Wang, A.R. Khan, N. Erkan, K. Okamoto, Effects of downward-facing surface type and inclination on critical heat flux during pool boiling, Ann. Nucl. Energy 113 (2018) 344–352. [37] S. Mori, Y. Utaka, Critical heat flux enhancement by surface modification in a saturated pool boiling: A review, Int. J. Heat Mass Transf. 108 (2017) 2534– 2557. [38] H.H. Son, N. Kim, S.J. Kim, Novel measurement of receding wicked liquid responsible for critical heat flux enhancement, Int. J. Heat Mass Transf 124 (2018) 150–157. [39] C. Paz, M. Conde, J. Porteiro, M. Concheiro, Effect of heating surface morphology on the size of bubbles during the subcooled flow boiling of water at low pressure, Int. J. Heat Mass Transfer. 89 (2015) 770–782. [40] C.Y. Lee, B.J. Zhang, K.J. Kim, Morphological change of plain and nano-porous surfaces during boiling and its effect on nucleate pool boiling heat transfer, Exp. Therm. Fluid Sci. 40 (2012) 150–158.
Contents lists available at ScienceDirect
International Journal of Heat and Mass Transfer journal homepage: www.elsevier.com/locate/ijhmt
A study on the effect of oxidation on critical heat flux in flow boiling with downward-faced carbon steel Kai Wang a,⇑, Nejdet Erkan a, Koji Okamoto b a b
Department of Nuclear Engineering and Management, School of Engineering, The University of Tokyo, 7-3-1 Hongo, Bunkyo-ku, Tokyo 113-8654, Japan Nuclear Professional School, School of Engineering, The University of Tokyo, 2-22 Shirakata, Tokai-mura, Ibaraki 319-1188, Japan
a r t i c l e
i n f o
Article history: Received 15 March 2019 Received in revised form 16 August 2019 Accepted 26 October 2019
Keywords: Critical heat flux Flow boiling Bubble behavior Heat transfer Oxidation effect
a b s t r a c t To execute in-vessel retention external reactor vessel cooling (IVR-EVRC) successfully, critical heat flux (CHF) plays a key role in securing the thermal and structural integrity of the reactor pressure vessel (RPV). In real-world applications, the RPV outer surface is exposed and oxidized. In this study, we performed a downward-face flow boiling CHF experiment using a carbon steel plate with several boiling cycles under water. Because the direction of gravity was inverse with the normal of the wall, it was difficult to remove bubbles, and in the local area, some local burning could repeat several times. After polishing the surface with sandpaper, the heat was supplied to the surface by cartridge heaters in a step-by-step manner until CHF was reached, after which the surface was cooled down. Using the same surface, without any further surface treatment, this heating cycle was repeated five times. The CHF value increased with the number of cycles. When the experiment was repeated for different mass fluxes, the CHF value also increased with the number of cycles. We suggest that this increasing tendency in CHF is associated with the increasing level of surface oxidation. After every repetition of the heating cycle, surface oxidation level increased, fewer bubbles were observed on the surface, resulting in higher CHF values. After the experiment, surface roughness increased and contact angle decreased. The liquidvapor mixture area decreased with increasing heat flux in one experiment, whereas it decreased with boiling time, under the same heat flux, which delayed CHF. The gradual oxidation process of carbon steel could be beneficial for real-world applications of IVR-ERVC. Ó 2019 Elsevier Ltd. All rights reserved.
1. Introduction In-vessel retention external reactor vessel cooling (IVR-EVRC) is a severe accident management strategy in which vessel breaching is prevented by flooding the reactor pressure vessel (RPV) cavity with water, so decay heat generated by the fuel rods can be removed by heat conduction through the RPV wall [1]. In this situation, critical heat flux (CHF) is likely to cause RPV failure by weakening the mechanical strength [2]. Consequently, CHF has been a common research focus in both pool and flow boiling experiments. Copper is widely used for pool boiling experiments, owing to its high thermal conductivity [3,4]. However, different materials have different CHF [5], and in real-world situations, lower plenum RPVs are composed of carbon steel. By performing pool boiling experiments with SA508, Kam et al. [6] found that as the width size decreased, the CHF enhancement ratio increased for carbon steel ⇑ Corresponding author at: 7-3-1 Hongo, Bunkyo-ku, Tokyo 113-8654, Japan. E-mail address: [email protected] (K. Wang). https://doi.org/10.1016/j.ijheatmasstransfer.2019.118966 0017-9310/Ó 2019 Elsevier Ltd. All rights reserved.
compared with stainless steel, especially at high inclination angles; they concluded that such factors as static contact angles and dissolved oxygen concentrations were not sufficient to explain the difference. Lee and Chang [7] conducted pool boiling CHF experiments with SA508 and SS304 test heaters under atmospheric pressure; they found higher CHF for SA508 and concluded that magnetite (Fe3O4) on the heater surface was produced by corrosion. Some studies have also been focused on flow boiling. Chang et al. [8] conducted flow boiling CHF experiments using a full scale 2D curved test section under atmospheric pressure with a SA508 Gr3 carbon steel heater; the results showed that SA508 displayed different (higher) CHF behavior in comparison with stainless steel, copper, and aluminum. Azzian et al. [9] tested carbon steel in a flow boiling loop designed specifically for IVR conditions. Up to 70% CHF enhancement values was observed for oxidized carbon steel in comparison with stainless steel. Park et al. [10] considered the effect of the heater material on CHF, finding higher CHF values for SA508 in comparison with stainless steel in all cases; they concluded that this behavior reflected the oxidation of carbon steel in the aqueous environment. Trojer et al. [11] showed that the
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pre-oxidized carbon steel had 70% CHF enhancement compared to that of the stainless steel in the flow boiling. Regarding other materials, Lee et al. [12] performed a saturated pool boiling experiment using zircaloy oxidized at 300, 450, and 600 °C, and increases in the CHF and wettability were found after the surface was oxidized. Lee et al. tried to explain the 40% CHF enhancement by the decreasing contact angle. Vlachou et al. [13] found that copper could also be oxidized and that this type of oxidation could decrease the heat transfer coefficient in the flow boiling. Coursey and Kim [14] obtained enhancements in both the CHF and wettability of copper after oxidation. Tachibana et al. [15] reported a CHF enhancement of oxidized aluminum CHF compared to that of the stainless steel. While most researchers have observed an increase in CHF with oxidation, Son et al. [16] found an inverse effect in which the CHF of bare SA508 was approximately 140% of Zuber’s pool boiling CHF prediction [17], whereas 3, 10, 20, 30, and 40 days of oxidation reduced the CHF of SA508 to approximately 65% of the Zuber’s prediction in the pool boiling.
Given the important role played by carbon steel in IVR, we investigated the CHF of carbon steel during downward face flow boiling instead of upward face boiling. The reason for conducting downward face boiling is that the lower head of RPV has a downward surface and the direction of gravity is inverse to the wall normal, so it is difficult for the bubbles generated on the lowest part to escape and CHF may occur earlier. In Wang et al. [18], a pool boiling experiment using carbon steel was carried out to investigate the CHF change due to local boiling. In the downward face boiling, it is very difficult for the bubbles to escape, so they may linger for a longer time. Thus, it is very likely that local burning may occur in some regions. This type of local boiling probably resulted from non-uniform heat flux around the lower head. The natural convection in the lower head pool caused by the temperature gradient between the pool boundary layer and the middle of the pool resulted in the non-uniform heat flux distribution.[19] In addition, the emission of fission gas and falling down of the melting fuels could cause the temperature and heat flux distribution to be unsteady. If at some point, the heat flux is greater
(a) Schematic of the experimental apparatus
(b) A photograph of the experimental apparatus Fig. 1. The experimental apparatus.
K. Wang et al. / International Journal of Heat and Mass Transfer 147 (2020) 118966
than the upper limit, CHF may happen. The heat flux may vary with time, making the heat flux of the local area change constantly. Thus, some points can experience multiple CHF situations. In Wang et al. [18] it was found that the CHF increased after several cycles in the pool boiling situations. However, few researchers have focused on the effect of flow boiling local burning on CHF. This type of local boiling may occur several times, so the local surface morphology can change through different cycles of local boiling. The boiling process and CHF will also change owing to the surface morphology change. Thus in this study, the local burnout of the downward face flow boiling was investigated.
2. Experimental setup and procedure 2.1. Experimental apparatus The schematic of the flow boiling experimental apparatus is shown in Fig. 1. The main components were the test section, three thermocouples in intervals of 3 mm for measuring the temperature of the copper block base, two high-speed cameras, one upstream tank with a thermocouple to measure the upstream temperature, and one downstream tank with a thermocouple to measure the downstream temperature. A pump with a maximum volume flow rate of 3 m3/h of distilled water was installed near the downstream tank. One electromagnetic flowmeter was applied to measure the
3
volume flow rate. Distilled water was supplied to the flow channel from the upstream tank. The flow channel was composed of acrylic transparent glass for visualization and had a cross section of 40 mm (width) by 10 mm (height). The test section was installed from the top of the channel to reach a downward heating surface. A reflux condenser was installed at the top of the downstream tank to condense overheated steam. Two identical high-speed cameras (Photron Fastcam SA5) were used. These cameras can operate with a maximum resolution of 1024 1024, 7000 fp. One camera was placed next to the water tank to capture images of the downward side reflected from a mirror under the water tank, and the other was placed near the water tank to obtain a side view. The two cameras were synchronized by connecting them together using the input/output function. A data logger connected to the five thermocouples was used for temperature measurement. The test section was composed of a copper base, carbon steel plate (heating zone), and polyether ether ketone (PEEK; thermal isolation zone) as shown in Fig. 2. The copper base was 40 40 mm rectangular on top and 30 30 mm on the bottom. The boiling surface of the carbon steel plate used here had an area of 30 30 mm rectangular. In total, nine cartridge heaters (each with a 225 W power rating) were inserted at the top of the copper block base controlled by a slidac. The reason that soldering method was adopted instead of using a whole carbon steel block was that the shielding material PEEK has a very low melting temperature, and the silicone used between
(a) Test section schematic
(b) A photograph of the test section Fig. 2. Test section schematic.
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the copper and PEEK cannot sustain temperatures greater than 200 °C. Moreover, the thermal conductivity of carbon steel is relatively low. High heat flux cannot be obtained by direct heating of the carbon steel block because the temperature would exceed the PEEK melting temperature; therefore, the carbon steel plate was attached to a copper using the silver soldering method. A similar method was also used by Mei et al. [20], who employed silver solder to adhere both carbon steel and stainless steel to the copper base. Detailed geometrical information on the copper block base is shown in Fig. 2. The carbon steel plate was approximately 1.0 mm. Golobicˇ and Bergles [21] found that when the heater surface exceeded a certain thickness, the CHF was stable. The asymptotic thickness is related to the thermal effusivity and for the carbon steel, and it was calculated to be approximately 0.36 mm, which is less than the thickness of the carbon steel plate (about 1.0 mm) used in this study. Thus, the effect of thickness could be eliminated in this study. After soldering, we measured the thermal conductivity of the silver solder layer. A thermocouple was attached to the bottom of the carbon steel surface and heat flux was applied by a slidac. Then, the temperature profile was recorded. The solder layer was measured to be approximately 0.1 mm and had a thermal conductivity of 15.0 W=ðm KÞ. The heat transfer through the heater block could be simplified using 1D steady-state heat conduction. The temperature of a carbon steel surface can be calculated by using Eq. (1), where T1 is the temperature of the copper base wall calculated using the 1D Fourier law,dcu ¼ 2 mm,kcu ¼ 400 W=ðm KÞ,dcs ¼ 1:0 mm, kcs ¼ 50 W=ðm KÞ [22], dsolder ¼ 0:1mm; ksolder ¼ 15:0 W=ðm KÞ:
T wall ¼ T 1 q
00
dcs dcu dsolder þ þ kcs kcu ksolder
ð1Þ
where dcu , dcs , and dsolder are the length of the copper, carbon steel, and silver solder layer, respectively, and kcu , kcs , and ksolder are the thermal conductivity of the copper, carbon steel, and solder layer, respectively. After oxidation, a layer of oxidized particles will be generated on the heating surface. The estimated layer thickness was 0.3 lm for a surface oxidized for 3 days [16], which is quite small compared to the thickness of the carbon steel plate. Thus, the oxidation effect on the temperature change of the wall was 00 not considered.q is the heat flux rate, which can be calculated using Eq. (2), and x denotes the distance between the copper thermocouples. In Jeong and Kim [23], the temperature gradient was also calculated based on the three thermocouples. They adopted a three-point, backward-space Taylor series approximation. In this study, the temperature gradient was calculated using a threepoint least-square fittings, as shown in Eq. (3). 00
q ¼ kcu
dT dx
ð2Þ
P P P dT ð3 xi T i xi T i Þ ¼ P 2 P dx 3 xi ð xi Þ2
ð3Þ
Here, xi denotes the distance from the copper base down face, and i = 1, 2, 3.T i is the temperature of the corresponding thermocouple. The total heat loss during the experiment was assessed using Eq. (4)
ahl ¼
V2 R
00
q S
ð4Þ
2
V R
00
where S is the area of the boiling surface, q is the heat flux derived from the thermocouple temperature difference, V is the voltage value measured by the multimeter connected to the slidac, and R is the resistance of the cartridge heater. The average heat loss percentage ahl was estimated to be approximately 4.9%, which includes both the heat dissipated to the air through the copper base or cartridge heater and the heat dissipated to the water through the silicon insulation layer.
2.2. Experimental procedure Owing to the easiness of carbon steel oxidation, the surface was completely cleaned prior to the experiment. Three types of sandpapers were used to clean the surface. To wipe off all residuals on the surface, very coarse P80 sandpaper was used firstly; then, fine P600 and P1200 sandpapers were used to polish the surface; and finally, the surface was cleaned by acetone. Before the experiment, we heated up the distilled water to its saturated temperature and the temperature was kept by controlling the heater. The air was evacuated for 30 min. It took approximately 2 h for heating and air evacuation. Therefore, before each experiment, the test piece had already been in the circulating water for approximately 2 h. During the experiment, the power of the cartridge heaters was increased gradually until CHF was reached. The voltage was increased by 20 V firstly, then 10 V, and 5 V when it came close to CHF, which was about 0.03 to 0.06 Mw/m2 of heat flux. The temperature increased and became stable at a certain point, and if the temperature was maintained for at least 1 min, we recorded this point and increased the heat flux for another point. CHF conditions were assumed when the temperatures of the thermocouples suddenly increased significantly (i.e., more than 50 °C over a very short time), and also at this time, a very long vapor film will form along the heating surface. The power supply was then stopped to prevent further damage. After CHF data were obtained and the power supply stopped, the temperature of both the water and block decreased. When the temperature of the block dropped to less than 95 °C, power was re-supplied until the next CHF value was
Table 1 Experimental uncertainties. Parameter
Measuring unit
Range
Uncertainty
Fluid temperature Inlet temperature Wall temperature Length Mass flux (160 kg/(m2 s)) Mass flux (320 kg/(m2 s)) Mass flux (640 kg/(m2 s)) CHF at 160 kg/(m2s) CHF at 320 kg/(m2s) CHF at 640 kg/(m2s)
K-type thermocouple N/A K-type thermocouple N/A Electromagnetic flowmeter Electromagnetic flowmeter Electromagnetic flowmeter
243.5–700 K N/A 243.5–700 K 0–3 mm 0–40 L/min 0–40 L/min 0–40 L/min
±0.75% 1% ±0.75% 3.3% 5% 2.5% 1.25% 6.1% 4.3% 3.7%
K. Wang et al. / International Journal of Heat and Mass Transfer 147 (2020) 118966
5
reached. Cycles were repeated five times for the same surface, without any surface cleaning between cycles. 2.3. Uncertainty analysis The estimated uncertainty of the heat flux was obtained using Eq. (5) [24]:
dR ¼ R
sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi du1 2 du2 2 dun 2 ð Þ þ ð Þ þ þ ð Þ u1 u2 un
ð5Þ
Table 2 Experimental flow rate parameters. Mass flux (kg/(m2 s)
Volume flow rate (L/min)
Velocity (m/s)
Reynolds number
160 320 640
4 8 16
0.17 0.33 0.67
0.91 104 1.82 104 3.64 104
(a) 160 kg/(m2 s)-A
Fig. 4. CHF enhancement with the boiling cycle for different mass fluxes.
(b) 160 kg/(m 2 s)-B
(c) 320 kg/(m2 s)-A
(d) 320 kg/(m2 s)-B
(e) 640 kg/(m2 s)-A
(f) 640 kg/(m2 s)-B
Fig. 3. Boiling curves for carbon steel plate. Experiment under a mass flux of (a) 160 kg/(m2 s)-A, (b) 160 kg/((m2 s))-B, (c) 320 kg/((m2 s))-A, (d) 320 kg/((m2 s))-B, (e) 640 kg/((m2 s))-A and (f) 640 kg/((m2 s))-B.
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Fig. 5. Pre- and post-experimental images of the carbon steel surface. (a) Pre-experimental surface. (b) Submerged in water. Experiment using a mass flux of (c) 160 kg/(m2 s), (d) 320 kg/((m2 s)), and (e) 640 kg/((m2 s)).
In this approach, the dependent variable R has independent linear parameters, such as R = R (u1, u2. . .un). The experimental uncertainties are reported in Table 1. The volume flow rate of the electromagnetic flowmeter is 0–40 L/min, and has ±0.2 L/min error range. This will result in a bigger uncertainty under smaller mass flux. 3. Results and discussion To investigate the enhancement effect of oxidation under different mass fluxes, three different mass fluxes were applied: 160, 320, and 640 kg/(m2 s). The relationship between volume flow rate V_ and mass flux G is
V_ ¼ GLH=ql Fig. 6. Average CHF as a function of mass flux. An empirical CHF correlation (Katto and Kurata [27]) is provided for comparison.
ð6Þ
where L (=40 mm) is the channel width, and H (=40 mm) is the channel height.ql (=958 kg/m3) is the density of water at saturation temperature.
Fig. 7. Images of the carbon steel surface during experiments under different mass fluxes and the same heat flux (=0.09 MW/m2). Experiments conducted at (a) 160 kg/(m2 s)1, (b) 160 kg/(m2 s)-2, (c) 160 kg/(m2 s)-3, (d) 160 kg/(m2 s)-4, (e) 160 kg/(m2 s)-5. Experiments conducted at (f) 320 kg/(m2 s)-1, (g) 320 kg/(m2 s)-2, (h) 320 kg/(m2 s)-3, (i) 320 kg/(m2 s)-4, and (j) 320 kg/(m2 s)-5. Experiments conducted at (k) 640 kg/(m2 s)-1, (l) 640 kg/(m2 s)-2, (m) 640 kg/(m2 s)-3, (n) 640 kg/(m2 s)-4, and (o) 640 kg/(m2 s)-5.
K. Wang et al. / International Journal of Heat and Mass Transfer 147 (2020) 118966
The relationship between velocity
v ¼ G=ql
v
and mass flux G is
ð7Þ
Reynolds number
Re ¼ GD=l
ð8Þ
Where l (=0.28 mPa s) is the viscosity of the water at saturation temperature, D (=16 mm) is the hydraulic diameter of the channel. Their velocities and Reynolds numbers are listed in Table 2. 3.1. Experimental boiling curves The boiling curves obtained in the experiments are shown in Fig. 3. For each experiment, the results from five CHF tests were obtained, and recorded as m kg/(m2 s)-n, where m is the mass flux in kg/(m2 s) and n is the cycle number. Experiments under different mass flux were conducted twice and was denoted as A and B. Generally, under low heat flux, the superheat increases with boiling cycle, indicating a deterioration of heat transfer. When the heat flux is close to CHF, the endpoint increases, confirming that the superheating and CHF increase with boiling time. The first CHFs of each mass flux were almost the same, indicating a good initial surface morphology control. However, after the first cycle,
7
the CHF enhancement became rather random due to the randomness of the oxidation process. In Wang et al. [18], experiments were done for 3 times at the same condition, in which the initial CHFs were the same in each experiment, but the CHF enhancement showed different ratios due to the randomness of the oxidation process. Although the experiments were done for a long time, it might not reach a saturate CHF. The reasons are as follows: After 5 cycles, the CHF has not reached a stable point. The oxidation process is a random process, and CHF enhancement is also different in each experiment. Fig. 4 shows the CHF enhancement of each experiment. In all the experiments, CHF showed different enhancement with boiling cycle. Images captured before and after each experiment show the condition of the surface (Fig. 5). The initial polished surface contained only a few scratches (Fig. 6a). After the experiment, the surface was immediately taken out from the water. Some surface oxidation occurred after all the experiments (Fig. 6c, 6d, and 6e). The surface roughness has been measured using a portable surface roughness tester SJ-210. After the test section was submerged in water for about 2 h, the test section showed a very slight increase in surface roughness as shown in Fig. 6b, which suggested that the surface received small amount of oxidation. After the experiment,
Fig. 8. Bubble film departure with a mass flux of 320 kg/(m2 s)-5 for different heat flux levels. Heat flux = 0.33 MW/m2: (a) down view, (b) side view; Heat flux = 0.81 MW/m2: (c) down view, (d) side view; post CHF (=0.87 MW/m2): (e) down view and (f) side view.
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the surface roughness was also measured and all the test surfaces had an increase in surface roughness. The increase of surface roughness was found in Son et al. [16] after oxidizing carbon steel in air for a long time. However, it was reported that the surface roughness had little effect on CHF. Hahne and Grigull [25] pointed out that the surface roughness might influence the vapor bubble growth but it became weaker near the peak heat flux. O’Hanley et al. [26] examined the separate effects of surface wettability, porosity, and roughness on the CHF using engineered surfaces, and concluded that surface roughness had no effect on CHF within the limit of the database. Several empirical and semi-empirical models have been proposed for CHF determination. Unfortunately, most of the correlations are based on vertical tube experiments. We chose an empirical CHF correlation (Katto and Kurata [27]) for comparison, which can be expressed as
qCHF ¼ 0:186Ghfg
qg qf
!0:559
rqf 0:264 G2 L
ð9Þ
whereG is the mass flux of the liquid, and L is the length of the heated surface. The correlation was developed based on vertical flow boiling experiments with a heating surface of 10 to 20 mm in length for saturated water and R-113. Comparison results are shown in Fig. 6. The first cycle CHF values all fell under the Katto and Kurata’s equation line; this is because the Katto and Kurata’s equation was derived from upward
flow boiling whereas our experiments were downward-face flow boiling. Downward-face flow boiling has a relatively smaller CHF owing to the difficulty of vapor removal because it was not easy for the bubbles generated on the downward face to escape from the surface due to gravity. An attempt was made to predict the CHF of downward face based on the Katto and Kurata’s equation. The modification was made mainly to the exponent of the term
rqf G2 L
, which is the ratio
of inertial force to surface tension force. The modification was done using a least-squares fitting method, and the results (Eq. (10)) showed good agreement with the results.
qCHF ¼ 0:235Ghfg
qg qf
!0:559
rqf 0:391 G2 L
ð10Þ
4. Discussion Detailed images of the downward surface were collected during the experiments (Fig. 7). It is observed that although not many, still a few bubbles came out from the heater sides. The bubbles generated on the edges showed some heat might be dissipated through the silicon layer between the PEEK and carbon steel plate, which was part of the heat loss explained in Eq. (4) in chapter 2. With the same heat flux, a longer oxidation time resulted in fewer but larger bubbles on the surface. The surface also became
Fig. 9. Surface images under different heat fluxes for an experiment with a mass flux of 160 kg/(m2 s)-5. Heat flux values of (a) 0.18 MW/m2, (b) 0.31 MW/m2, (c) 0.47 MW/m2, and (d) 0.68 MW/m2.
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progressively darker, indicating increased level of oxidation of the carbon steel during the boiling process. Greater oxidation was consistent with fewer but larger bubbles. Studies on the impact of the nucleation site density (NSD) on the heat transfer coefficient and CHF have yielded various results. For increasing CHF, Jaikumar et al. [28] fabricated the microstructure and performed a pool boiling experiment, which showed that both the heat transfer coefficient and CHF increased due to the change of NSD and wettability. Kurihara and Myers [29] reported that the nucleate boiling enhancement was due to an increase in active NSD. Betz et al. [30] found that hydrophobic surfaces can delay CHF. Chang and You [31] showed that that increasing active nucleation sites increased pool boiling heat transfer coefficients by 30%, and
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increases CHF by approximately 100%. On the contrary, You et al. [32] found that nanoparticles in water decrease the number of bubbles and increased the CHF. Yeom et al. [33] reported a reduction in the NSD and an increase in the CHF in the pool boiling experiment using nanoparticle coatings formed by electrophoretic deposition of Ti and TiO2 particles. Wang et al. [34] found an inverse relationship between NSD and CHF by irradiating a copper block. They concluded that greater NSD generates more bubbles and increases the heat transfer coefficient; furthermore, the area of bubble vapor covering the surface increases, which prevents water supply to the boiling surface and results in the early occurrence of CHF. Wang et al. [18] also reported that the CHF was enhanced with oxidation and fewer bubbles generated on the sur-
Fig. 10. Surface images during subsequent heating cycles under similar heat flux and a mass flux of 160 kg/(m2 s). Image of (a) cycle 1, 0.37 MW/m2, (b) cycle 2, 0.33 MW/m2, (c) cycle 3, 0.32 MW/m2, (d) cycle 4, 0.32 MW/m2, and (e) cycle 5, 0.31 MW/m2.
Fig. 11. Contact angle of the carbon steel surface. (a) Before experiment, (b) Submerged in water. After the experiment under a mass flux of (c) 160 kg/(m2 s), (d) 320 kg/(m2 s), (e) 640 kg/(m2 s).
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face in the pool boiling. Wang et al. [35] observed both an increase and a decrease of CHF due to the change in NSD. They identified a threshold NSD (<~5 105 sites/m2), above which the CHF was deteriorated due to the increase in the NSD and vice versa. Our experimental results are consistent with those of a pool boiling experiment conducted by Wang et al. [18], in that a decrease in NSD delayed the CHF. This delay probably occurred because in the downward face boiling, it is easy for bubbles to accumulate due to the difficulty of bubble removal, so more bubbles will delay CHF occurrence. We found that under relatively small heat flux, bubbles larger than a certain size moved and left the surface, but under medium–high heat flux, small bubbles coalesced to form a bubble film. Using images captured by high-speed cameras, we could easily distinguish that under medium heat flux, this bubble film was able to move off the surface. Fig. 8 shows bubble film departure for different heat fluxes. According to the side view camera, most of the bubble film generated on the surface did not exceed the channel height, which meant that channel height would not restrict the bubble behavior. Actually, Katto and Kurata [27] used a rectangular nozzle with a width of 15 mm, and two heights 10 mm mainly and 5 mm auxiliary. And they found that difference of CHF was hardly perceived for the change of channel height from 10 mm to 5 mm. For a medium heat flux (Fig. 8a and b), the bubble film departs the surface without connecting to the next bubble film, which means that little vapor was produced. For high heat flux (Fig. 8c and d), the departing bubble film has some connections to neighboring bubble films, which means some water was prevented from supplying to the surface. Fig. 8e and f show CHF images; the bubble film extends continuously across the surface, showing that as the heat flux increases, bubble vapor continues and no water can be supplied for heat removal. Fig. 9 shows surface images under different heat fluxes with a mass flux of 160 kg/(m2 s)-5. Initially, several small bubbles coexisted with liquid (i.e., the liquid–vapor mixture area); however, as heat flux increased, this area decreased and more vapor was generated to cover the surface more, preventing water from reaching to the surface and thus leading to CHF. Fig. 10 shows surface images of each experimental cycle with a mass flux of 160 kg/(m2 s) and a constant heat flux. There is an increase in the mixture area with time, resulting in the delay of CHF appearance, which probably reflects a decrease in the NSD. This phenomenon was also observed by Gong et al. [36], who used a copper surface and found that after irradiation, the liquid-vapor mixture area was enlarged compared with the non-irradiated surface. They attributed this enlargement to a change in hydrophilicity after irradiation. A decrease in the contact angle (i.e., an increase in the wettability) after oxidation has been observed in various studies [6,12,16], and was also found in our research (Fig. 11). The contact angle was measured before and after the experiment. It showed clearly that the contact angle decreased after the experiment. Mori and Utaka [37] noted that CHF enhancement may be due to a change of NSD, wettability and capillary wicking. Lee et al. [12] explained the 40% CHF enhancement of oxidized zircaloy by the increasing wettability. However, the contact angle alone cannot explain the CHF enhancement, as recently reported by Son et al. [38], who stated that nanostructure, nanoporous, and microstructure surfaces exhibited CHF enhancement compared to a bare surface, as well as smaller contact angles and increased liquid wicking. In this experiment, it is difficult to quantify the wicking effect due to the non-uniform surface change. Nevertheless, the change in contact angle and surface morphology indicated that the surface property change did influence flow boiling, but it was difficult to identify whether it was the wicking effect.
To confirm these surface changes, energy dispersive spectroscopy (EDS) of the oxidized surface was performed. Due to the difficulty of conducting EDS experiments on the surface after experiment, small round test pieces heating in the water under saturation temperature to simulate the actual experimental environment were observed instead. Fig. 12 showed the EDS results of the surface which were non-oxidized, 1 h oxidized in water and 3 h oxidized in water, respectively. For all the test specimens, the peak of iron, oxygen and carbon were detected. The mass content of oxygen increased from 1.09%, 2.90% to 5.90%, which clearly showed iron oxidation. In the non-oxidized specimen, very small amount of oxygen was detected, which might be some oxygen in air environment. The oxygen content was also detected in Chi et al. [12], which the non-oxidized zircaloy surface also contained small amount of oxygen. The non-oxidized and oxidized carbon steel surfaces were also imaged by backscattered electron imaging using an SEM (Fig. 13). Scratches and small cavities are evident on the surface without oxidation in Fig. 13a and 13b, and some oxidation particles had
(a) Original test piece
(b) 1 h oxidized test piece
(c) 3 h oxidized test piece Fig. 12. Surface EDS of (a) original test piece, (b) 1 h oxidized test piece and (c) 3 h oxidized test piece.
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Fig. 13. SEM images of the non-oxidized and oxidation carbon steel. Images acquired at (a) 4000 magnification for non-oxidized carbon steel, (b) 10000 magnification for non-oxidized carbon steel, (c) 1000 magnification for oxidized carbon steel, and (d) 8000 magnification for oxidized carbon steel.
Fig. 14. Schematic of the gradual oxidation of a carbon steel surface.
occurred within these scratches, covering the cavities and decreasing the NSD in Fig. 13c and 13d. This mechanism for the gradual oxidation of the carbon steel surface is illustrated in Fig. 14. In fact, the filling of the pores and cavities has also been reported. For example, Paz et al. [39] carried out subcooled flow boiling experiments using copper test section with different roughness, and noted that the surface morphology should be a very important factor but has usually been ignored in correlations and experiments. Lee et al. [40] also noted that the decrease in the heat transfer coefficient was due to the surface morphology change
caused by the formation of aluminum hydroxide on the boiling surfaces. This process causes the pores between the nanostructures to be filled and smoothed.
5. Conclusions In this study, flow boiling CHF experiments were conducted on a downward facing surface under different mass fluxes to investigate the effect of oxidation. The main conclusions are as follows:
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(1) During the boiling process, the CHF increased with boiling cycle gradually as fewer bubbles generated on the heating surface. (2) After the experiment, surface roughness became greater than the original test surface. (3) The correlation based on the Katto and Kurata’s equation for downward-face agreed well with the experimental results. (4) The liquid-vapor mixture area decreased with increasing heat flux in the experiment. Furthermore, with increasing boiling time, the mixture area decreased under a constant heat flux, which delayed the CHF. (5) The increasing CHF contributed to increasing the wettability and decreasing NSD. In real-world applications, RPV surfaces tend to corrode with time, and the changing surface characteristics impact the boiling phenomena and CHF significantly. In the real RPV lower head, because the direction of gravity is inverse with the normal of the wall, bubbles are difficult to be removed and in some area, some local burning may repeat several times. Our results demonstrate the advantages of using EVRC to dissipate the residual heat that the CHF increases with the local burnout. However, additional research is necessary to examine whether this type of enhancement will continue to increase or saturate with further cycle repetition. Declaration of Competing Interest We declare that we have no conflict of interest. Acknowledgements This work is mainly supported by the Japan Grant-in-Aid for Scientific Research (B) (Grand No. 19H02645). The author would like to thank Dr. Huilong Yang from Nuclear Professional School, School of Engineering, The University of Tokyo for his help on the SEM and EDS analyses. References [1] J.L. Rempe, K.Y. Suh, F.B. Cheung, S.B. Kim, In-vessel retention of molten corium: lessons learned and outstanding issues, Nucl. Technol. 161 (3) (2008) 210–267. [2] T.G. Theofanous, C. Liu, S. Additon, S. Angelini, O. Kymaelaeinen, T. Salmassi, In-vessel coolability and retention of a core melt, Nucl. Eng. Des. 169 (1–3) (1997) 1–48. [3] M.S. El-Genk, Z. Guo, Transient boiling from inclined and downward-facing surfaces in a saturated pool, Int. J. Refrig. 16 (6) (1993) 414–422. [4] M.S. El-Genk, A.F. Ali, Enhanced nucleate boiling on copper micro-porous surfaces, Int. J. Multiph. Flow. 36 (10) (2010) 780–792. [5] K. Wang, N. Erkan, H. Gong, L. Wang, K. Okamoto, Comparison of pool boiling CHF of a polished copper block and carbon steel block on a declined slope, J. Nucl. Sci. Technol. 9 (55) (2018) 1–14. [6] D.H. Kam, Y.J. Choi, Y.H. Jeong, Critical heat flux on downward-facing carbon steel flat plates under atmospheric condition, Exp. Therm. Fluid Sci. 90 (2018) 22–27. [7] J. Lee, S.H. Chang, An experimental study on CHF in pool boiling system with SA508 test heater under atmospheric pressure, Nucl. Eng. Des. 250 (2012) 720–724. [8] H. Chang, T. Hu, W. Lu, S. Yang, X. Zhang, Experimental study on CHF using a full scale 2-D curved test section with additives and SA508 heater for IVR-ERVC strategy, Exp. Therm. Fluid Sci. 84 (2017) 1–9. [9] R. Azizian, T. McKrell, K. Atkhen, J. Buongiorno, Effects of porous superhydrophilic surfaces on flow boiling critical heat flux in IVR accident scenarios, American Nuclear Society, Chicago, 2015. [10] H.M. Park, Y.H. Jeong, S. Heo, Effect of heater material and coolant additives on CHF for a downward facing curved surface, Nucl. Eng. Des. 278 (2014) 344– 351.
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