Cement and Concrete Research 45 (2013) 69–78
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Lubrication layer properties during concrete pumping Myoungsung Choi a, c,⁎, Nicolas Roussel b, Youngjin Kim a, Jinkeun Kim c a b c
Civil Engineering Research Team, Daewoo Institute of Construction Technology, Suwon, Republic of Korea Université Paris Est, IFSTTAR, Paris, France Department of Civil and Environmental Engineering, Korea Advanced Institute of Science and Technology, Daejeon, Republic of Korea
a r t i c l e
i n f o
Article history: Received 5 January 2012 Accepted 22 November 2012 Keywords: Fresh concrete (A) Rheology (A) Modeling (E) Pumping
a b s t r a c t In order to progress in the understanding of the physical phenomena involved in the pumping process of concrete, we study in this work the properties of the lubrication layer, which forms between the pipe and the bulk material. Using ultrasonic velocity profiler, we measure the thickness of this lubrication layer in the case of real size pumping circuits. Our experimental, analytical and numerical results suggest that, from a rheological point of view, this layer behaves similarly as the constitutive mortar of the pumped concrete. Moreover, this layer thickness does not seem to depend on flow rate. We finally propose a simple analytical relation allowing for a rough estimation of the pumping pressure. © 2012 Elsevier Ltd. All rights reserved.
1. Introduction Concrete pumping is extensively used worldwide for the construction of high rise buildings and long span bridges. This technique offers notable advantages such as allowing concrete casting in difficult to access locations, and reducing casting process durations or allowing for a continuous concrete casting. However, in order to apply concrete pumping to large-scale construction projects, a priori prediction of concrete flow rates, which induce the construction process duration, is necessary from a practical logistic point of view. This prediction shall be based on the analysis of the concrete flow typology within a pumping pipe. This typology is complex but it is now accepted that flow of concrete in a pipe strongly differs from the one of typical viscous fluids such as water or oil. The primary reason for this difference comes from the fact that concrete is a yield stress fluid (i.e. it flows only if the applied stress is higher than its yield stress). As a consequence, there exists at the center of the pipe (i.e. around the symmetry axis where the shear stress is equal to zero) a zone where concrete is not sheared [1]. Most approaches of concrete pumping in literature have taken into account this yield stress and the existence of this unsheared zone [2–4] by assuming that concrete behaves as a Bingham fluid or Herschel Buckley fluid. These approaches have however almost systematically failed to predict correctly pumping flow rates on large range of concrete fluidity without including the fact that there exists a lubrication layer at the interface between concrete and the pipe. ⁎ Corresponding author at: Civil Engineering Research Team, Daewoo Institute of Construction Technology, Suwon, Republic of Korea. Tel.: +82 10 5604 1095; fax: +82 32 250 1148. E-mail address:
[email protected] (M. Choi). 0008-8846/$ – see front matter © 2012 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.cemconres.2012.11.001
Indeed, the second reason for the difference between pumping of concrete and pumping of simpler materials comes from the fact that, under the action of shear, a redistribution of particles occurs within the pipe. This is a common feature of particle suspensions and initially well mixed particles in concentrated suspension flows are shown to undergo migration from high shear rate regions to low shear rate regions [5–7]. For example, in a Couette viscometer with a rotating inner cylinder and a stationary outer cylinder, the particles migrate towards the outer cylinder [8] whereas in Poiseuille flows in cylindrical pipes, the particles migrate towards the central axis [9]. This particle migration, in the case of coarse particles (i.e. particles with a characteristic size close to the characteristic size of the flow), can be increased by wall effect at the interface between the pumped material and the pipe [10]. Indeed, because of simple geometrical considerations, it is not possible to find the center of a particle of diameter a at a distance from a wall lower than a/2. During pumping, shear concentrates therefore in a fluid layer of material depleted from the coarsest particles of the concrete. In the inner region, the material is almost not sheared (Cf. Fig. 1). Pumping of concrete can therefore be considered in most cases as the shearing of an annular layer of unknown thickness and made of a material with unknown rheological properties. This layer is often called in literature lubrication layer or sometimes slippage layer. Its existence was first suggested by Aleekseev [11] and Weber [12]. Morinaga [13] noted that, from a theoretical point of view, considering the rheological parameters of the material, pumping of concrete would not be possible without the formation of this lubrication layer. Sakuta et al. [14] went further and showed that the flow properties of the bulk material were irrelevant. The only property that matters is the ability of the material to form this layer. Jacobsen et al. [1] conducted experimental research with colored fresh concrete flowing after ordinary concrete to observe the flow conditions in
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Fig. 1. Schematic pattern of concrete flow in the pipe.
various pipes and Rossig [15] pumped some colored concretes in a pipe for a direct observation of flow profiles. Their results demonstrated the existence of a high velocity and paste rich zone at the vicinity of the pipe wall. From a quantitative point of view, the thickness of this lubrication layer was estimated to be between 1 and 5 mm in [2,3,16]. Feys et al. [4] noted that the thickness and rheological properties of this layer seem to depend on the mix proportions of the pumped concrete. The macroscopic consequences of this layer on the pumping pressure were taken into account by introducing in the pumping process prediction some interface properties measured macroscopically using a tribometer by [3]. Considering the critical role played by this lubrication layer on concrete pumping and considering that a detailed analysis of this layer thickness and properties has not been carried out yet, we study in this work the properties of this layer. Using ultrasonic velocity profiler, we measure the thickness of this lubrication layer in the case of a real size pumping circuits equipped with pressure gauges. Our experimental, analytical and numerical results suggest that, from a rheological point of view, this layer behaves similarly as the constitutive mortar of the pumped concrete. Moreover, its thickness does not seem to depend on flow rate. We finally propose a simple analytical relation allowing for a rough estimation of the pumping pressure.
2. Materials and protocols 2.1. Materials and mix proportions Three different highly-workable concretes (i.e. fluid concretes) were studied in this work. Their mix proportions are given in Table 1. The cement was a CEM I 52.5 N with density of 3150 kg/m3. The sand was a natural river sand with a density of 2590 kg/m 3 and fineness modulus of 2.81. Sand particle size ranges from 0.08 to 5 mm with water absorption capacity of 2.43%. The maximum coarse aggregate size was 25 mm. It was a limestone aggregate with water absorption capacity of 0.8%, density of 2610 kg/m 3, and fineness modulus of 6.72. The amount of mixing water was corrected to take into account the water absorbed by sand and coarse aggregates. A polycarboxylate-based high range water reducing admixtures (HRWRA) was used. Its dosage is shown
Table 1 Mix proportions. Materials
Design strength
Name of the series
C40
C50
C60
Cement CEM I 52.5 N, kg/m3 Fly ash Class F, kg/m3 Blast furnace slag, kg/m3 W/B ratio Sand, kg/m3 Coarse aggregate, kg/m3 % Polycarboxylate-based HRWRA Slump flow, mm
201 45 201 0.38 768 873 0.8 600 ± 20
225 50 225 0.33 736 871 0.9 620 ± 20
257 57 257 0.28 713 844 1.0 620 ± 20
Fig. 2. Overview of the experimental setup.
in Table 1 marked as % HRWRA, meaning percentage of admixture relative to binder content (in weight). Each concrete was produced in a batch of 2 m3 by a ready mix company. The mixing procedure was as follow: sand and coarse aggregates were mixed during 15 s; all other dry components (cement, filler) were added during 15 s; water and HRWRA were added during the additional 2 min of mixing. The total mixing time was 2 and half minutes. 2.2. Pumping circuit A horizontal pumping circuit of 170 m length was set up (Cf. Fig. 2). This circuit included eight 180° and three 90° bends. The pipe diameter was 125 mm and its thickness was 7.7 mm. The concrete pump was a high pressure piston pump (model BSA2110HP-D, Putzmeister Co., Germany). Its specifications are given in Table 2. For high pressure pumping capacity, a piston side cylinder type was used. Three different flow rates were studied. The filling rate of the pump cylinder, which directly affects flow rate, was calibrated from specific experiments prior to the actual concrete pumping. The concrete was pumped into several 1 m3 reservoirs, which were connected to a Linear Variable Differential Transformer (LVDT). As the pump cylinder volume is known, the filling rate was computed from the LVDT length variation with designated stroke times. Through these experiments, the averaged filling rate was found to be around 85%. Based on this value, the stroke times needed to target the three flow rates were chosen as shown in Table 3. The circuit was equipped with 11 pressure gauges (model JUMO dTRANs p20, JUMO GmbH & Co. KG, Germany). The detailed locations of the gauges are shown in Fig. 3. The first one was located 5 m after the beginning of the circuit whereas the last one was located 11 m before the end. 2.3. Ultrasonic velocity profiler (UVP) An Ultrasonic Velocity Profiler (UVP) [17] was used to get a non-destructive measurement of the concrete velocity profile in the pipe using ultrasonic waves with high resolution. The detailed specifications of the device used in this work are given in Table 4. UVP allows for Table 2 Pump specifications. Item
Content
Model Flow rate (m3/h) Max. pressure (bar) Engine horsepower (kW) Stroke/min
BSA2110HP-D 102⁎/76⁎⁎ 150⁎/220⁎⁎ 330 28⁎/19⁎⁎
* Rod side, ** piston side.
M.S. Choi et al. / Cement and Concrete Research 45 (2013) 69–78 Table 3 Pump setup.
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Table 4 Experimental UVP parameters.
No
Stroke time (s)
Average velocity (m/s)
Flow rate (m3/h)
Item
Content
1 2 3
4 5 7
1.13 0.92 0.68
50 40 30
Frequency (MHz) Cycles per pulse No. of profiles Sound velocity (m/s) Doppler angle (°) Spatial resolution (mm)
8 2 ~ 32 1024 2680 ± 200 85 ± 0.5 Min. 0.20
the measurement of the flow velocity profile through the following steps: pulsed emission of ultrasound signals, echo reception and detection of the Doppler shift frequency. Experimental investigations of pipe flows using ultrasonic waves have already been carried out in the literature [18–21] for the measurement of velocity profiles of simple fluids such as water or dilute suspensions. However, applications to concentrated suspensions of coarse particles such as concrete are far more limited. In order to apply this device to measure the concrete flow in the pipe, a stable positioning of the ultrasonic probe (transducer) is essential for the successful measurement as precision of measurement cannot be higher than precision of transducer positioning. This especially concerns the angle of transducer. Any deviation of the ultrasound beam axis from the mechanical axis is at the origin of high measurement uncertainties. The position of the effective ultrasound beam axis was therefore checked by installing a wire across the ultrasound beam, while simultaneously observing the echo from the wire on an oscilloscope. As the maximum amplitude of ultrasound is located on the beam effective axis, this procedure allows for the identification of the location of the effective beam axis. An overall schematic illustration of this device is shown in Fig. 4. In addition, for application of the UVP using ultrasonic waves, 1 m transparent engineered plastic pipe with the same diameter as the standard pipe was installed in the last section, as shown in Fig. 5. When ultrasonic waves propagate in a medium containing coarse particles such as concrete, the ultrasound pulse hits the particle and part of the ultrasound energy is scattered and lost for the echo measurement. Thus, as the measuring depth increases, the amplitude of echoed ultrasound energy decreases. This is especially true for the high frequency ultrasonic waves used in this work. Thus, above a given depth, the amplitude of echoed ultrasound energy is not sufficiently strong to detect the Doppler shift frequency and give access to the velocity profile. In this work, with the chosen test condition, only the velocity profile within 15 mm from the wall can be measured. Although this limited measurement range does not give access to the full velocity profile in
the cross section, it exceeds the expected thickness of the lubrication layer.
2.4. Rheological measurements In order to measure the rheological properties of the constitutive mortar, a Brookfield DV-II viscometer was used [22]. The container is filled up with 0.5 L of mortar. The radii of the inner spindle and container are 8 mm and 36 mm, respectively. The height of the spindle is 60 mm. The range of rotational speed is 0.4 to 4.2 rev/s, which corresponds to approximately 60 s −1 maximum shear rate. The rheological properties of concretes were measured with a coaxial cylinder type ConTec Viscometer 5 [23]. The radii of the inner cylinder and outer cylinder are 100 mm and 145 mm, respectively. The height of the inner cylinder is 98 mm. Both the inner and outer cylinders are equipped with ribs to prevent slippage between the tested concrete and the steel surface. The range of rotational speed is 0.1 to 0.6 rev/s, which corresponds to approximately 10 s−1 maximum shear rate. The constitutive mortar and concrete were tested at the age of 15 min after water addition time. In each viscometer, the measuring procedure was initiated with a 30 s high speed phase to eliminate any thixotropy and structural breakdown artifacts [24,25]. The rotational velocity was then decreased stepwise. Between two rotational velocity steps, we had a 2-second transient time and 3-second sampling time to reach what was defined as steady state. A Bingham model was then fitted to the data allowing for the computation of plastic viscosity and yield stress. More information and details on the measuring and data transformation procedures for the Brookfield DV-II and ConTec Viscometer 5 can be found in [26–28]. Because of the existing discrepancy between concrete rheometers [29,30], it can be expected that the absolute values of the rheological parameters obtained on concrete could be debatable. Moreover, shear induced particle migration could also occur in the rheometer.
Fig. 3. Schematic ground plan of the pumping circuit and the location of pressure gauges.
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Ultrasonic probe Particles moving in measuring volume
Flow direction
Reflection from bottom Ultrasound echo generated by particles
time Ultrasound burst emission time
Velocity
Fig. 4. Schematic illustration of the UVP velocity profile measurement on a flow.
However, as reminded in the introduction and as discussed further in this paper, the contribution of the bulk to the pumping pressure can be neglected compared to the contribution of the lubrication layer as long as a lubrication layer is formed. Orders of magnitudes of the bulk rheology are therefore sufficient within the frame of this paper. 3. Experimental results and analysis 3.1. Pressure measurements The measured pressure distribution along the pipe is illustrated in Fig. 6, in which the pressure waveforms measured at each location for
C50 are shown for the three tested flow rates. The averaged 0 m position values of pressure waveforms (i.e. without the strokes) which are calculated by using linear extrapolation process of 11 designated position pressure gauges are shown in Fig. 7 for the three tested concretes as a function of flow rate. The measured pressures show an almost linear relationship with flow rate regardless of mix proportions with an extrapolated ordinate at the origin in the studied regime almost equal to zero. This indicates that, in this regime and in the pumping conditions tested in this work, the pumping pressure does not seem to be affected by a yield stress (i.e. a nonlinear material behavior), by shear thickening or shear thinning [31–35] or by any pressure dependency of the rheological parameters of the pumped materials [36,37]. In addition, if we consider now the pressure gauges #3, #4, #9 and #10 in Fig. 3, which are located before and after a bend, our measurements suggest that the presence of the bend does not affect the local pressure drop. This result is similar to the one reported in Kaplan et al. [3] but differs from, for example, Schwing [38] which proposed that a bend of 90° causes a pressure loss which is equivalent to 3 m of straight pipe.
3.2. UVP measurements
Fig. 5. Application point of the UVP and transparent engineered plastic.
The axial velocities measured by using UVP are shown in Fig. 8. A brutal change in slope can be spotted for all concretes tested and at all flow rates. It can be seen in these figures that shear rate (i.e. approximately the slope of the velocity profiles) concentrate in a layer, the thickness of which does not seem to depend on flow rate and mix
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a) Flow rate: 50 m3/h (stroke time 4 s)
b) Flow rate: 40 m3/h (stroke time 5 s)
c) Flow rate: 30 m3/h (stroke time 7 s)
Fig. 6. Pressure waveforms for the 11 pressure gauges along the pipe for C50.
proportion in the ranges studied in this paper. The value of this thickness is approximately 2 mm. It is moreover interesting to note that, after a parabolic increase of the velocity in the lubrication layer with shear rates of the order of a
a) C40
few hundreds s −1, a low shear rate of the bulk can be measured (around 10 s −1). No plug flow (i.e. zero shear rate) can be spotted. It has however to keep in mind that only the first 15 mm of flowing material are considered here.
b) C50
Fig. 7. Averaged pressure values as a function of flow rate for the three tested concretes.
c) C60
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a) C40
b) C50
c) C60
Fig. 8. UVP measurement results for the lubrication layer thickness and its variations depending on the mix proportion and flow rate.
3.3. Material properties As discussed in Section 1, particle collisions in highly sheared and/or highly concentrated zones force particles to migrate from these zones. This effect is counterbalanced by the local increase in concrete viscosity resulting from this migration (i.e. in the bulk). Shear induced particle migration finds therefore its origin in the competition between gradients in particle collision frequency and gradients in viscosity of the concrete within the pipe. It was shown in [39] that the characteristic time for shear induced particle migration is of the order of R2 =10a2 ϕ2 γ_ where R is the pipe radius, a is the particle diameter, ϕ is the particle volume fraction and γ_ is the shear rate. As a consequence, sand particles, the average diameter of which is one order of magnitude lower than the average diameter of the gravel shall migrate 100 times slower. This means that, when sand particles migrate, they encounter the high viscosity of the bulk concrete, in which gravel particles have already migrated and should be prevented to migrate inside the bulk. It can therefore be expected that the migration of sand particles can be neglected compared to the migration of gravel particles and that the lubrication layer could be considered, as a first approximation, as being similar to the constitutive mortar of the pumped concrete. We chose here to test this assumption and wet-screened the constitutive mortar from the fresh concrete and measure its rheology in the Brookfield viscometer. We simultaneously measure the rheology of the concrete with ConTec Viscometer 5 (Cf. Section 2.4). The results are gathered in Table 5. These rheology parameters are the averaged values before pumping.
and the pipe. We choose therefore in the following to develop a very basic and simple approach focusing specifically on this lubrication layer and not on the flow of the concrete bulk. As described below, this approach is however limited to the case of fluid concretes (i.e. low yield stresses). Because of the momentum conservation, the shear stress in the pipe as a function of the radial coordinate writes τ = ΔP ⋅ r/2L, where ΔP/L is the pressure gradient, which is constant along the circuit as shown in Fig. 6. As the lubrication layer thickness δ is small compared to the radius of the pipe, the average shear stress τl in the lubrication layer can be approximated as τl = ΔP ⋅ R/2L. The results from the previous sections have shown that the shear rate in the lubrication layer is of the order of several hundreds s −1. In this shear rate range, the contribution of the yield stress of the mortar can be neglected and the average shear rate γ_ l in the lubrication layer can be approximated as γ_ l ¼ ΔP⋅R=2Lμ pl , where μpl is the plastic viscosity of the mortar in the lubrication layer. Similarly, in the bulk, the average shear stress τb and the average shear rate γ_ b can be respectively approximated as τb = ΔP ⋅ R/4L and γ_ b ¼ ΔP⋅R=4Lμ pb , where μpb is the plastic viscosity of concrete in the bulk. This last relation can only apply if the shear rate in the bulk is sufficient to neglect the contribution of the yield stress. This assumption is valid as long as τ0b =μ pb γ_ b bb1, where τ0b is the yield stress of
4. Analytical approach and numerical simulations 4.1. Analytical approach As stated in Section 1, most analytical approaches in the literature have almost systematically failed to predict correctly pumping flow rates on large range of concrete fluidity without including the fact that there exists a lubrication layer at the interface between concrete
Table 5 Rheology parameters. Type
C40
Item
Lubrication Bulk Lubrication Bulk layer layer
Plastic viscosity (Pa∙s) 2.0 Yield stress (Pa) 5.0
C50
30.0 70.0
2.5 10.0
C60 Lubrication Bulk layer
40.0 3.0 100.0 15.0
60.0 180.0
Fig. 9. Predicted shear rates in the lubrication layer and in the bulk as a function of the measured shear rates for the three tested concretes.
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Fig. 10. Pumping pressures as a function of flow rate for the three tested concretes. The dashed lines are the computed values from the analytical approach.
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Fig. 12. Pressure variation as a function of flow rate for various lubrication layer thicknesses.
the bulk concrete. This is almost the case for the concretes and flow rates tested in this work as the above ratio is always between 0.2 and 0.4. As a consequence, the ratio between γ_ l and γ_ b writes γ_ b =γ_ l ¼ μ pl =2μ pb . This means that the shear rate in the lubrication layer is between 30 and 40 times higher than the shear rate in the bulk. This value is in agreement with our UVP measurements (Cf. Fig. 8). The average shear rate in the pipe can be approximated as V0/R, where V0 is the average concrete velocity given in Table 3. It results from the contribution of both γ_ l and γ_ b . The average shear rate in the lubrication layer can therefore be estimated as: V0 γ_ l ¼ δ þ Rμ pl =2μ pb
ð1Þ
We compare in Fig. 9 the computed and measured values of the shear rates in the lubrication layer and in the bulk for the three concretes and three flow rates tested. We conclude that the above simple approach seems to be able to capture the concentration of shear in the fluid lubrication layer. This correct agreement moreover suggests that the rheology of the material in the lubrication layer is the one of the constitutive mortar. With the approximated average shear rate, the pressure gradient can be computed as: 2μ pl V 0 ΔP ¼ L R δ þ Rμ pl =2μ pb
ð2Þ
The pumping pressures computed from the pressure gradient are plotted in Fig. 10 as a function of the measured flow rate for the three tested concretes. Although several strong simplifications were
a) Overall mesh for 170 m conduits
Fig. 13. Pressure variation with respect to the length considering the lubrication layer effect.
made, there is a good agreement between the computed and measured pressures. However, as mentioned above, there exist several limits to a practical application of Eq. (2). First, as already stated above, it shall only apply to low yield stress concretes. The average shear in the bulk could be reduced in the case of high yield stress concretes leading to a higher average shear rate in the lubrication layer, which cannot be predicted by Eq. (2). Second, the thickness of the lubrication layer is a very important input parameter, which is not easy to access. Although it is
b) Cross section including the 2 mm lubrication layer (dark region)
Fig. 11. Modeling for numerical simulations.
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constant for all concretes tested in this paper and equal to 2 mm, it can be expected that it could vary with the mix design of the pumped concretes and potentially with the pipe diameter. Shear induced particles migration reaches equilibrium because of the increase in viscosity of the zones where particles are migrating to. It can then be expected that, as suggested in [39], shear induced particle migration shall stop
when local particle volume fraction is high enough to reach the so-called random loose packing ϕRLP (i.e. the lowest volume fraction value allowing for a percolated network of contacts between particles). It was shown recently that this random loose packing can be estimated from and is proportional to the value of the maximum packing fraction ϕm of the particles [40,41]. At the entrance of the pumping circuit, the
a) C40
(a-1) Flow rate: 30 m3/h
(a-2) Flow rate: 40 m3/h
(a-3) Flow rate: 50 m3/h
(b-2) Flow rate: 40 m3/h
(b-3) Flow rate: 50 m3/h
(c-2) Flow rate: 40 m3/h
(c-3) Flow rate: 50 m3/h
b) C50
(b-1) Flow rate: 30 m3/h
c) C60
(c-1) Flow rate: 30 m3/h
Fig. 14. Comparison of pressure distribution along the pumping circuit between numerical simulation results and experimental data.
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bulk is composed of both sand particles and gravel particles at volume fractions equal to the mix design volume fractions. Shear induces a migration of the coarse particles from the outer lubrication layer until the volume fraction of sand and gravel in the bulk reaches ϕRLP. Steady state is then reached and the thickness of the lubrication layer does not change any more. The fact that pumpable concretes are often mix designed with lower amount of coarse particles than traditional concretes allows therefore for the lubrication layer formation. It can be naturally expected that the extent of coarse particle migration and therefore the lubrication layer thickness shall increase with a reduction in the initial volume fraction of gravel. This question cannot however be settled in the case of the three concretes tested here as they all have almost the same initial volume fractions of sand and gravel with similar proportions of each. This leads to similar values of ϕm and ϕRLP and similar potential of coarse particle migration. Finally, it can be kept in mind that the above frame suggests that the lubrication layer thickness at steady state does not depend on flow rate as steady state equilibrium between particle collisions in the highly sheared lubrication layer and the local increase in concrete viscosity in the bulk shall only depend on the sand and gravel particle initial volume fractions and therefore on concrete mix design.
4.2. Computational fluid dynamics approach In the above analytical approach, we neglected the yield stress contribution in both lubrication layer and bulk concrete. This assumption is valid in the lubrication layer but could be debatable in the bulk. In order to predict accurately flow rates as a function of the all measured rheological parameters and compare them to the experimental values, computational fluid dynamics (CFD) were therefore used to solve this complex flow. The computational modeling techniques found in the literature and able to simulate concrete flow may be divided into three categories [42,43]: single phase fluid, particles suspended in a fluid and discrete particle flow. In this work, the single phase fluid method was chosen and performed with the commercial CFD code Fluent [44] and the following boundary conditions were used: The inlet velocity was fixed at the average velocity computed from flow rate (Cf. Table 3). No slip conditions were chosen for the pipe interface. At the end of the pumping circuit, the pressure was fixed at the atmospheric pressure. The computational zone was divided into a lubrication layer and a bulk zone (Cf. Fig. 11). In each zone, the corresponding rheological properties were given to each fluid material. As a first step, we considered the case of a 170 m pumping circuit with a 50 m 3/h flow rate and varied the thickness of the lubrication layer from 1 to 5 mm (Cf. Fig. 12). The best agreement was obtained for a layer thickness of 2 mm as measured by the UVP. As a second step, we compared the pressure profile along the pumping circuit with and without a 2 mm lubrication layer for the concrete C50. As shown in Fig. 13, the numerical simulation results considering the lubrication layer coincide fairly well with the experimental data. If the lubrication layer is not considered, the predicted pumping pressures are approximately three times higher than those of the experimental data. It can moreover be noted that the pumping pressure required to pump the concrete almost exceeds the capacity of the high pressure pump. This would be the case of a concrete unable to form the lubrication layer (i.e. high coarse particles mix design volume fraction). Finally, with the 2 mm lubrication layer thickness (CF Fig. 11(b)) and rheological properties from Table 5, numerical simulations of the 170 m pumping circuit studied in this work were performed and compared with experimental data. The pressure distribution along the pumping circuit for the three tested concretes and the three flow rates are shown in Fig. 14. The discrepancies between calculated results and experimental data are lower than 7%.
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5. Concluding remarks In order to progress in the understanding of the physical phenomena involved in the pumping process of concrete, we studied in this work the properties of the lubrication layer, which forms between the pipe interface and the bulk material. Using ultrasonic velocity profiler, the measured thickness of the lubrication layer was found to be around 2 mm for the concretes, pumping systems and flow rates tested here. This layer thickness does not therefore seem to depend on flow rate but shall depend on the sand and gravel particle initial volume fractions and therefore on concrete mix design and potentially on pipe diameter. Our experimental, analytical and numerical results all indicate that, from a rheological point of view, the lubrication layer behaves similarly as the constitutive mortar of the pumped concrete. In order to predict accurately flow rates for all measured rheological parameters and compare them with the experimental values, computational fluid dynamics (CFD) were performed considering the lubrication layer properties. These simulations were able to predict accurately the obtained measurements. We moreover proposed a simple analytical relation allowing for a rough estimation of the pumping pressure and obtained a good agreement with the measured pressures. It can be finally noted that, based on the above results, it seems possible to estimate pumping pressure from the rheology of both wet screened mortar and concrete, thickness of the lubrication layer and geometry of the pumping circuit. Finally, it can be kept in mind that the thickness of the lubrication layer shall depend on the concrete mix design and potentially on pipe diameter. Further research on this correlation is needed.
Acknowledgment The authors would like to thank the colleagues at DICT for their assistance, in particular Yongjic Kim, and Junhee Jo. They also, appreciate the valuable comments and review from Prof. Seung-Hee Kwon at Myongji University and Prof. Se-Jin Jeon at Ajou University. This research was supported by a grant from the Construction Technology Innovation Program (08CTIPE01-Super Long Span Bridge R&D Center) funded by Ministry of Land, Transportation and Maritime Affairs (MLTM) of Korean government.
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