M M Salem

E De Luycker

M Fazzini

P Ouagne

Experimental, analytical and numerical investigation to prevent the tow buckling defect during fabric forming ☆

We investigate the causes, kinematic and possible ways to prevent the tow-buckling defect during the complex shape-forming of a dry woven reinforcement. Macro-scale compression leads to wrinkles; we focus on a mesoscale phenomenon generated by a non-uniform axial compression of the tow due to in-plane bending. An experimental study is presented, followed by an efficient method to predict the onset of the tow-buckling defect. This investigation is based on the combination of an experimentally validated analytical approach, with a macroscale simulation of the forming process for a single fabric layer. A particular focus is given to the resulting tow curvatures that allows us to predict the zones of high tow buckling probability. The relatively simple numerical approach based on Lagrange assumptions does not take into account tow slippage. However, for most cases, where buckles appear before slippage, this approach alleviates the compulsory need for an expensive meso-scale representation.

Introduction

Fiber reinforced structural composites offer great mechanical performances because of their low density compared to regular structural materials. Manufacturing such parts requires the use of a fibrous reinforcement. Among others, woven fabrics represent a large amount of the chosen reinforcement due to their high drapability, good permeability and good cohesion when subjected to multidirectional loading and high manufacturing processability [START_REF] Greb | Fabrics for reinforcement of engineering composites[END_REF]. However, defects that may jeopardize the mechanical integrity of the composites could appear, prior to resin injection, during the complex shape forming of woven reinforcements such as the resin transfer molding (RTM) process, as identified by Potter et al. [START_REF] Potter | Variability, fibre waviness and misalignment in the determination of the properties of composite materials and structures[END_REF]. Some of these defects, such as the preform wrinkling, have been extensively investigated in the literature. Several experimental investigations [START_REF] Lightfoot | Defects in woven preforms: formation mechanisms and the effects of laminate design and layup protocol[END_REF][START_REF] Nezami | Analyses of interaction mechanisms during forming of multilayer carbon woven fabrics for composite applications[END_REF][START_REF] Rashidi | Passive control of wrinkles in woven fabric preforms using a geometrical modification of blank holders[END_REF] and numerical investigations [START_REF] Boisse | Simulation of wrinkling during textile composite reinforcement forming. Influence of tensile, in-plane shear and bending stiffnesses[END_REF][START_REF] Hosseini | Identifying the distinct shear wrinkling behavior of woven composite preforms under bias extension and picture frame tests[END_REF][START_REF] Hübner | Simulation-based investigations on the drape behavior of 3D woven fabrics made of commingled yarns[END_REF] led to the understanding of the mechanics behind the formation of the wrinkling defect and the ways to prevent it (blank holders design, ply orientations, etc.).

Other defects mentioned by Potter et al. [START_REF] Potter | Variability, fibre waviness and misalignment in the determination of the properties of composite materials and structures[END_REF] remain to be fully understood, such as, tow sliding/unweaving as experienced in [START_REF] Nezami | Analyses of interaction mechanisms during forming of multilayer carbon woven fabrics for composite applications[END_REF][START_REF] Allaoui | Experimental preforming of highly double curved shapes with a case corner using an interlock reinforcement[END_REF][START_REF] Ouagne | Complex shape forming of a flax woven fabric; analysis of the tow buckling and misalignment defect[END_REF] and studied in [START_REF] Gatouillat | Meso modelling for composite preform shaping -simulation of the loss of cohesion of the woven fibre network[END_REF][START_REF] Labanieh | Intra-ply yarn sliding defect in hemisphere preforming of a woven preform[END_REF] or the tow buckling defect.

The tow buckling defect was defined by Ouagne et al. [START_REF] Ouagne | Complex shape forming of a flax woven fabric; analysis of the tow buckling and misalignment defect[END_REF][START_REF] Ouagne | Analysis of defects during the preforming of a woven flax reinforcement[END_REF] as a tow deflection out of the plane of the reinforcement by a few millimeters combined with a rotation around its axis. Capelle et al. [START_REF] Capelle | Complex shape forming of flax woven fabrics: design of specific blank-holder shapes to prevent defects[END_REF] designed a special set of blank holders associated with a tetrahedral punch to limit the appearance of tow buckles. Tephany et al. [START_REF] Tephany | Development of an experimental bench to reproduce the tow buckling defect appearing during the complex shape forming of structural flax based woven composite reinforcements[END_REF] and Ouagne et al. [START_REF] Ouagne | Measurement of the appearance and growth of tow buckling defect in the frame of complex shape manufacturing process by using fringe projection technique[END_REF] found from their preliminary study the causes that could lead and influence the tow buckles appearance. The said causes could be linked to the previously mentioned ways to prevent the wrinkling defect. A device was designed to recreate the conditions of the tow buckling defect and investigate the causes. They concluded, that the most influencing parameters are the in-plane bending of the tow, the reinforcement orientations and the applied loads in both warp and weft directions of the reinforcement. However, some other geometric and mechanical parameters were not investigated. Beakou [START_REF] Beakou | Modelling slit tape buckling during automated prepreg manufacturing: a local approach[END_REF] and Matveev [START_REF] Matveev | Understanding the buckling behaviour of steered tows in Automated Dry Fibre Placement (ADFP)[END_REF] described a comparable phenomenon on carbon tape during the automatic layup process. They deducted that the lateral buckles are a result of the dissymmetrical in-plane longitudinal compression load generated in the tape via in-plane bending; they proposed an analytical solution to predict its appearance. This work sets itself as a follow-up of Tephany's work [START_REF] Tephany | Development of an experimental bench to reproduce the tow buckling defect appearing during the complex shape forming of structural flax based woven composite reinforcements[END_REF] and a more in-depth investigation of the effects of tensile loads, nature and geometrical parameters of the tows on the tow buckling defect. Our aim is to better understand the defect appearance and the possibilities to prevent it. In conjunction with an experimental study, the analytical via each of the jaws bolts. The lateral jaws can then follow an in-plane circular translation motion as illustrated by Fig. 1(b,d).

In addition to the two load sensors that allow tension measurements, the device is instrumented with two Stereo Digital Image Correlation systems (S-DIC), as shown in Fig. 2, associated with the model proposed in [START_REF] Beakou | Modelling slit tape buckling during automated prepreg manufacturing: a local approach[END_REF][START_REF] Matveev | Understanding the buckling behaviour of steered tows in Automated Dry Fibre Placement (ADFP)[END_REF] was modified so that to predict the appearance of tow buckling in the woven fabric reinforcements. This paper is divided in two parts: the first one dealing with experimental investigation and the second one presenting the analytical modeling and the possibility to predict the tow buckling appearance during the forming of a complex shape, such as a prismatic shape, using an isogeometrical simulation. The aim of our numerical work is in the same vein as the work of Iwata et al. [START_REF] Iwata | Coupled meso-macro simulation of woven fabric local deformation during draping[END_REF]; predicting defects appearance. The chosen numerical approach is much simpler though with the known limitations of macro-scale modeling which are unable to consider tow slippage contrary to meso-scale modeling. An isogeometrical framework is chosen for its capability to naturally evaluate the curvatures along the tows during a simulation. The main advantage here is the computational efficiency of the macro-scale description since no contact is computed within the fabric. Only contacts between the fabric and the tools are considered. It is therefore possible to perform such simulations on a laptop computer and it permits investigating a possible prediction of the tow buckling defect.

Experimental study

Materials and experimental setup

For this study, the same device (Fig. 1) as the one described in [START_REF] Iwata | Coupled meso-macro simulation of woven fabric local deformation during draping[END_REF][START_REF] Wu | Buckling analysis and optimisation of variable angle tow composite plates[END_REF] was used and modified. The device permits to impose different initial tensile loads in warp and weft directions of a woven fabric as represented in Fig. 1(a,c). The loads can therefore be applied in the fixed longitudinal tow network (direction 1, load 1) and in the transversal tows network submitted to in-plane bending (direction 2, load 2) software GOM ARAMIS®. The first set of S-DIC set retrieves global information at the scale of the whole device (Fig. 3(a)) and the second one focuses on the formation of the tow buckles locally (Fig. 3(b)). The in-plane bending state of the reinforcement is evaluated in two ways. A global way via a so-called "bending angle of the device" defined in Tephany [START_REF] Tephany | Development of an experimental bench to reproduce the tow buckling defect appearing during the complex shape forming of structural flax based woven composite reinforcements[END_REF], as presented in Figs. In order to evaluate the buckling behavior of a tow within a fabric, at each bending state, its most deflecting profile is extracted from the 3D reconstructed upper surface of the fabric (Fig. 4(a)). The different states of the same profile are then compiled in Fig. 4(b). The red curve in Fig. 4(b) represents the final state of the profile of the buckled tow at a bending angle of ~40°.

In order to determine the causes for the tow buckling defect, previously stated parameters such as tensile loading and bending angle were investigated more in depth alongside other fabric geometrical parameters such as the tow free dimensions and the tows nature. The tow free dimensions are the unsupported (free-floating) length and width of the apparent portion of the tow that is expected to buckle as illustrated in Fig. 3(b). Several reinforcements, with similar tow widths and/or unsupported lengths, were considered. Reinforcements such as flax fabrics, manufactured specially in the frame of a previous project by Groupe Depestele (France), and a carbon fabric (HexForce 48,600 U 1250) manufactured by Hexcel (France) and which characteristics are listed in Table 1.

Experimental results

Effect of the tows dimensions and fabric construction

To determine the effect of the tow geometrical properties, reinforcements manufactured with the same tows but different weaves were selected. This allows us to vary the tow unsupported length and width independently from other parameters. The selected reinforcements were made from desized-up flax tows and subjected to an initial load 1 of 300 N and a load 2 of 20 N. Each reinforcement is bent in its plane up to a bending angle of ~40°. From the experiments, the tow buckles height and tow rotation (Fig. 5) are extracted. The buckle height is calculated by subtracting the maximal height of the initial profile of the tow H 0 max from the maximal height of the deflected profile H max . The tow rotation is calculated from the angle between the chord of the initial profile and the chord of the buckled profile. The influence of the unsupported dimensions on the tow buckles height and tow rotation during buckling is presented in Fig. 6(a,b). Fig. 6(a) shows that the maximal height and rotation of the tows increase together with the unsupported length. This is explained by the increase in the mobility of the buckling tow allowing it more room to rise and rotate. It causes a higher perceivable buckle. Fig. 6(b) shows that the increase of tow width still increases the maximal registered height whereas the rotation level decreases. Each of the transversal buckling tow is supported underneath it by a longitudinal fixed tow. As the tow buckles upward, it does so by combining two motions, i.e., deflection and rotation. During the early rotation stages of the buckle, the inner edge lifts and the outer edge has a tendency to sink until the whole tow is fully deflected. The narrower the tow, the faster the whole tow lifts completely from the supporting tow for lower bending angles. Fig. 6(a,b) also show that a higher bending angle is required to fully lift a larger buckling tow from its supporting tow. For the same final bending angle, wider tows have less rotation with a slight increase in height (60% rotation angle loss against 6% height increase for a 42% larger tow) as illustrated in Fig. 7. Consequently, the tow buckles look smaller.

Effect of the tow type

In order to study the effect of tow type on the buckling, three different reinforcements manufactured from three different tows but with similar woven structures were chosen so that to keep similar dimensions, i.e. sized-up flax, desized-up flax and carbon fiber all in 2 × 2 twill weaves. Each tow of the studied fabrics have different tensile, bending and torsional rigidities that are expected to influence the buckling mechanism. However it is difficult to experimentally decouple each of those rigidities. Consequently, we studied the tow type as a whole. The properties of the tows are presented in Table 2 with the methods used to obtain them explained in the parameters identification Section 3.2.

During the tests, each reinforcement is loaded in tension with 300 N in load 1 and 20 N in load 2. This load is spread over 20 tows. Each reinforcement is bent in-plane up to an angle of ~40°. The tow maximal height and the tow rotation are represented respectively in Fig. 8(a) and (b). Fig. 8(a) shows that the buckles elevation is higher for the carbon tows than the other considered tows. The elevation of the buckle also begins from lower in-plane bending angles (~8°). One can also see that the sized-up flax tows exhibit higher maximal tow elevation than the desized-up ones.

A direct correlation between the increase of tow rigidities and the buckle elevation is observed. It is also possible to observe that for the three reinforcements considered, the appearance of the tow buckles happens for lower in-plane bending angles when the tow rigidities increases.

Fig. 8(b) also shows that the rotation of the buckle happens from low in-plane bending angle and the tow rotation angle is much larger for the most rigid carbon tows than the two flax tows. In the case of the two flax tows, the tow rotation angle evolution is similar which is probably due to the relatively low difference in their rigidities. This correlates well with the observation of the maximum elevation of the tow buckle.

Effect of the tows tension in direction 1: fixed tow network (load 1)

To study the effect of load 1 on the buckles, a series of tests has been conducted on 5 samples of 2 × 2 twill weave fabrics manufactured from sized-up flax and carbon tows respectively while imposing each time an initial load in direction 1 between 20 N and 400 N spread over 20 tows. The load in direction 2 is initially 20 N. Each sample is bent in-plane up to an angle of ~40°. The results of the test series are given in Fig. 9(a) for the sized-up flax fabric and (b) for the carbon reinforcement.

As shown in both Fig. 9(a) and (b), the maximal height reached by the tow evolves oppositely to the curvature radius of the tow. The lower the final curvature radius (increase in the tow curvature) and the higher the tow buckles. The final rotation angle, on the other hand seems to be fairly constant and independent of the final curvature radius. When considering the evolution of both elevation and tow rotation, Fig. 10 gives an idea on the kinematics involved during the bucking. On early bending stages, the height is mainly due to the rotation of the tow; then, as the tow continue to bend in its plane, the rotation seems to stabilize while the elevation further increases. This in turns suggests that the elevation at this point is mainly due to a global deflection of the tow that can be seen in the different tows final profiles in Fig. 11. The reason behind the parabolic trend of the curvature radius (Fig. 9(a,b)) can be attributed to the combination of the in-plane bending imposed by the Table 1 Characteristics of the different reinforcements used.

jaws alongside the tow sliding (up to 2.81 mm) as represented in Fig. 12.

During the applied in-plane bending on the device, tow sliding may take place as shown in Fig. 12 and this phenomenon may have consequences on the final curvature. The tow sliding as explained by Labanieh et al. [START_REF] Ouagne | Measurement of the appearance and growth of tow buckling defect in the frame of complex shape manufacturing process by using fringe projection technique[END_REF] depends on the contact pressure between the warp and the weft tows. As the load 1 increases, the tow sliding is inhibited going from 2.8 mm at a load 1 of 20 N to 1.1 mm at a load 1 of 400 N for the sized-up flax. This explains the three stages presented in Fig. 9(a,b), which are described further below:

-For load 1 from 20 N to 200 N: a low pressure is applied on the perpendicular tows cross-points. This allows the tows expected to buckle to slide, preventing them from achieving their potential full curvature. -At ~200 N: an optimum is reached. The in-plane bending and tow sliding leads to the maximal possible curvature of the tows submitted to in-plane bending. -Above 200 N: the pressure applied on the cross-points combined with the friction phenomenon prevent the tow from sliding and from curving leading to a decreased curvature compared to the optimum. When considering the influence of the increasing load 1 on the tow buckling within the flax and carbon-based fabrics, similar behavior was observed. It indicates that keeping low loads in direction 1 may delay the appearance of the buckles and reduce their height since tow sliding taking place (in a limited extent without the appearance of gaps). This sliding tend to reduce the curvature of the tow submitted to in-plane bending. These results confirm that a reduction of the blank-holder pressure as performed by Capelle et al. [START_REF] Matveev | Understanding the buckling behaviour of steered tows in Automated Dry Fibre Placement (ADFP)[END_REF] on the tows perpendicular to the ones exhibiting the defect permits to delay the appearance of the buckles. On the other hand, when the load in direction 1 is larger than 200 N for both the fabrics, the high contact pressures prevent the tow sliding but also limits the in-plane bending of the tows. Rising the load in direction 1 to high tension values may therefore limit the curvature and consequently delay the appearance and reduce the height of the buckles. However, a risk of gaps due to tow sliding may happen, and such approach may be considered with a great care.

Effect of the tows tension in direction 2: tow network submitted to inplane bending (load 2)

As for load 1, the study was done on sized-up flax with only the initial load 2 as a variable for each of the samples going from 20 N to 400 N. The results are summarized in Fig. 13.

As observed with load 1, the maximal height registered evolves oppositely to the curvature radius of the tow while the rotation angle is similar for all loads. The decrease of the curvature is explained by an increase in load 2 that adds more slide to the tows and prevents them from curving or buckling. Increasing the load in direction 2 should therefore favor a reduction of the buckles size.

Analytical analysis

In order to be able to predict the buckling onset during the forming of a woven structure, an analytical model is set up. This approach was previously used in the literature by Beakou et al. [START_REF] Beakou | Modelling slit tape buckling during automated prepreg manufacturing: a local approach[END_REF] and Matveev et al. [START_REF] Matveev | Understanding the buckling behaviour of steered tows in Automated Dry Fibre Placement (ADFP)[END_REF] in order to address the problem of tape buckling during the automated fiber placement following a curved path. The following work is inspired by their analytical framework.

Analytical model

Fibrous materials exhibit complex behavior; due to the possibility for fibers to slide; cross sections during bending do not necessarily remains perpendicular to the neutral fiber. The Kirchhoff-Love plate theory (and corresponding Euler-Bernoulli beam theory) are most of the time not representing well the behavior of fabrics, interlocks, and virgin fiber bundles. The bending rigidity doesn't relate to the tension one. This is illustrated by the difference of bending behavior between a phone book with the pages glued together (tension/compression is generated) and without glue (bending of each page taken separately plus friction). In this part of the work, fibers inside a flat tow, in small strain (only the onset of buckling is considered) possess cohesion, due to the entanglement of fibers and due to sizing (for desized-up flax, we assume that a small amount of sizing remains). With this assumption, we consider that fibers don't slide inside a tow and that Kirchhoff plate theory is valid; a link can then be made between the kinematic (the curvature) and the load (axial compression P) as illustrated Fig. 14.

If the tow is considered as a thin homogenous orthotropic plate [START_REF] Beakou | Modelling slit tape buckling during automated prepreg manufacturing: a local approach[END_REF][START_REF] Matveev | Understanding the buckling behaviour of steered tows in Automated Dry Fibre Placement (ADFP)[END_REF], simply supported at its outer edges and clamped on its radial edges by the neighboring tows, the differential equation for plate bending, in Eq. ( 1), can be solved by using the Rayleigh-Ritz approach [20]:

∇ = w P 2 ( 1
)
where w is the plate deflection, which, alongside its first derivative w', are null on the clamped edges of the tow, represented in Fig. 15, (w (x = 0) = w(x = L) = 0 and w'(x = 0) = w'(x = L) = 0). The outer edge is supposed to be simply supported and only its deflection w is null (w(y = 0) = 0). Under those assumptions, the deflection w can be described as a linear combination of the n th first modes and approximated, in Eq. ( 2), based on the boundary conditions and the compressive load formulated previously in Eq. ( 1). A closed-form solution in a Rayleigh-Ritz approach can be considered:

∑ = ⎛ ⎝ -⎞ ⎠ ⎛ ⎝ - ⎛ ⎝ ⎞ ⎠ ⎞ ⎠ = + w x y c y b πmx L ( , ) 1 1 cos 2 i n i i 1 1 ( 2
)
where m is the buckling mode. Only the first one (m = 1) will be taken into account because it corresponds to the most critical buckling mode. P the compressive load, caused by the in-plane bending of the tow is given by Eq. ( 3):

= ⎛ ⎝ - ⎞ ⎠ P y P αy b ( ) 1 0 (3)
where P 0 is the minimal buckling load applied at the inner edge, as represented in Fig. 14, of the bent plate and α is the load non-uniformity coefficient also represented in Fig. 14. The Rayleigh-Ritz approach dictates that the total energy Π, represented in Eq. ( 4), should be minimized by differentiating it with respect to the deflection's amplitude, c i , as shown in Eq. ( 5):

= - w U w Q w Π( ) ( ) ( ) (4)
= d dc Π 0 i ( 5
)
where U(w) is the elastic strain energy of the plate which is defined in Eq. ( 6) and Q(w) is the work of the compressive load as defined in Eq. ( 7):

∫ ∫ ⎜ ⎟ ⎜ ⎟ ⎜ ⎟ ⎜ ⎟ ⎜ ⎟ = ⎛ ⎝ ⎜ ⎛ ⎝ ⎞ ⎠ + ⎛ ⎝ ⎞ ⎠ + ⎛ ⎝ ⎞ ⎠ ⎛ ⎝ ⎞ ⎠ + ⎛ ⎝ ⎞ ⎠ ⎞ ⎠ ⎟

U w d w dx d w dy d w dx d w dy d w d dydx

( ) 1 2 D11 D22 2D12 4D66 xdy L b 0 0 2 2 2 2 2 2 2 2 2 2 2 2 (6) ∫ ∫ ⎜ ⎟ = ⎛ ⎝ - ⎞ ⎠ ⎛ ⎝ ⎞ ⎠ αy b d w dx dydx Q(w) 1 2 P 1 L b 0 0 0 2 2 2 ( 7
)
where L is the length of the plate, b the width of the plate and Dxx are the bending stiffness of the plate. Replacing w for (n = 1) and in Eqs. (2), ( 6) and ( 7) and then minimizing Π with respect to c 1 , the critical buckling compressive load, Pcr is obtained as given by Eq. ( 8): During bending, taking into consideration small deformations and the intra-tow cohesion of fibers, the edge of the tow is assumed to bend following the curvature of a radius R at an angle θ according to Kirchhoff's plate theory, represented in Fig. 14. The length of the inner side of the tow caused by the bending is found to be equal to θ(R-b/α) and the deformation of the tow is equal to b/(αR). The load in the inner edge, P 0 , generated by in-plane bending of the tow can be written [START_REF] Kubiak | Static and dynamic buckling of thin-walled plate structures[END_REF][START_REF] Balcioǧlu | An investigation on lateral buckling of laminated composites with delamination[END_REF] as:

= P E hb αR 1 0 (9)
In the case where the minimal Pcr (that is calculated for a reinforcement values of L: tow unsupported length, h: tow thickness and b: unsupported tow width) is equal to P 0 , we are able to connect the critical load to the geometrical bending parameters of the tow and in particular to find the critical buckling radius Rcr as: that can be recovered from the laminate plate theory, where, the bending stiffness matrix for a single layer [START_REF] Kubiak | Static and dynamic buckling of thin-walled plate structures[END_REF][START_REF] Balcioǧlu | An investigation on lateral buckling of laminated composites with delamination[END_REF] For a dry tow with minimalistic interactions in the transverse direction between the constituting fibers (mainly introduced by the false twisting during yarn manufacturing) compared to the interactions alongside the fibers in the longitudinal direction. The cohesion in the transversal direction is expected to be much smaller than in the longitudinal direction. This results, for the bending, in B x ≫ B y and with L > b we can safely assume that D11 ≫ D22 and therefore, D22 can be neglected. Furthermore, to the knowledge of the authors, no experimental procedures exists to this day to characterize the transversal bending behavior of a tow and this assumption can not be concluded upon.

= - + + - ( ) Rcr E α b h α 1 (6
D12 and D21 are equal and represent the bending coupling, which, by Poisson action effect, is the cause of an anticlastic deformation (saddle effect). Considering that the tow is fairly narrow [START_REF] Swanson | Anticlastic effects and the transition from narrow to wide behavior in orthotropic beams[END_REF] and the beveled sides of the tow [START_REF] Conway | Anticlastic curvature of strips of variable thickness[END_REF], it is safe to assume that their effect has a negligible influence on the buckling compared to the other stiffnesses. D66, the twisting bending stiffness or torsional stiffness is the torque required to generate a twist K z :

= D G h 66 12 6 3 (14)
With G12 being the shear modulus. The critical buckling radius, Rcr, from Eq. ( 10) can then be written as follows:

= - + Rcr E α bh L α D L D π b m 1(6 ) (160 66 24 11 ) 3 2 2 2 22 ( 15
)
This critical buckling radius is expected to correspond to the radius at which the tow buckling appears, but several parameters need to be experimentally evaluated.

Parameters identification

In order to calculate the critical buckling radius, the following parameters E1, α, D11 and D66 need to be identified. The following sections are dedicated to the experimental identification of the parameters.

Tensile modulus E1 identification

The tensile modulus E1, was determined using a tensile test on each of the considered tows. The slope of the linear part of the stress-strain curve was calculated and reported in Table 2.

Bending stiffness D11 identification

To identify the bending stiffness, D11, in Eq. ( 12), we must determine the bending rigidity, B x . It is possible to do so experimentally using the Peirce cantilever test [START_REF] Peirce | The "Handle" of cloth as measurable quantity[END_REF] where the tow is bent under its own mass as described in Fig. 14. The bending rigidity B x can calculated using Eq. ( 16):

= ( ) B L M g θ cos 8tan( ) x l θ 3 2 (16)
where L is the bent length of the tow, M l is the linear density, g is the gravity and θ is defined in Fig. 16(b). This means that Eq. ( 12) can be written with all the parameters that are constant or measurable as shown in Eq. ( 16):

= ( ) D L M g θ b 11 cos 8tan( ) l θ 3 2 (17)
Following the experimental investigation, the results are presented in Table 2 for four different tows taken from different woven fabrics.

Torsional stiffness D66 identification

To evaluate D66 from Eq. ( 14), it is necessary to evaluate the inplane shear modulus, G12, which is calculated from a torsion test that can be performed on an ARES rheometer as shown in Fig. 17(a) and where the torsional rigidity is given by Eq. ( 18) [26]:

= TL φ G Jt 12 ( 18
)
where T is the torque, L is the length of tow subjected to rotation, φ is the torsional angle, J t is the polar moment of inertia (similar to second moment of the area for bending) that for a rectangular surface is given by Eq. ( 19):

⎜ ⎟ ⎜ ⎟ = ⎛ ⎝ - ⎛ ⎝ - ⎞ ⎠ ⎞ ⎠ J bh h b h b 1 3 0.21 1 12 t 3 4 4 (19)
T L/φ can be obtained from a torsion test by averaging the 0 to 2π and 0 to -2π slopes of the torque T as a function of the unit angle φ/L curve as represented in Fig. 17(b). The tow is subjected to an initial tensile loading up to 100 N. Finally, D66 from Eqs. ( 14) and ( 18) can be written as:

= D T L h φ Jt 66 6 3 (20)
The experimental results for the bending stiffness D66 are also reported in Table 2.

Non-uniformity coefficient α identification

The non-uniformity coefficient was obtained from observation on the tow using digital image correlation using the GOM ARAMIS® software on images such as the one presented in Fig. 18. Zones circled in red are areas representative of tensile loading while zones circled in blue represent compression loading areas. Both areas seem to be fairly equal. The hypothesis of pure bending can be assumed and thus α = 2 was chosen according to the representation given in Fig. 16.

Calculation of the critical buckling radius

The theoretical critical radii were then calculated and compared to the experimental critical radii (tow in plane bending at which buckles appear) in Table 2. As there is no clear definition in the literature for what constitutes the threshold of the buckle's onset, we opted to determine the experimental critical radii by the first sudden increase in the derivative of the tow elevation dz with respect to the bending radius R (e.g. dotted line in Fig. 19 for plain weave C), represented in Fig. 19. The aforementioned Figure should be read from highest curvature radius to the lowest (right to left).

The difference between the experimental and the calculated critical radii are given in Table 2. One can observe that the differences are relatively low with an average error of 8.8%. The fact that the critical radii obtained from the experimental and the analytical approaches are close indicate that the analytical approach permits to predict with a relatively good accuracy (and a global tendency to underestimate) the critical bending radius at which the tow buckles are expected to appear. The observed differences can be due to simplification hypothesis such as the homogeneity assumption of the tow and uncertainties in the radius measurement.

As the analytical model globally permits to predict well the in-plane curvature radius at which the tow buckles are expected to appear, this method could be coupled to a simulation of the textile forming capable to predict the curvature of the tows with a good accuracy. An example of the use of such a combined approach is described in the next section of this work.

Modelling of the defect appearance in a macroscopic simulation

Once the critical bending radius is determined experimentally or analytically, this result can be of great interest in a macroscopic simulation of the forming process of dry fabrics. Important effort has been paid in the past 20 years in order to predict the mechanical response of fabrics during forming at macroscopic scale using numerical simulations. The reached maturity allows now to confidently predict the shear angles via the determination of yarn local orientations. It's then straightforward to use those yarns orientation to evaluate, with a high level of confidence, the curvature of the yarns within a fabric structure during forming. Comparing the local in-plane curvature of the tow (determined during the macroscopic simulation of the textile forming) to the critical one, we can evaluate locally the risk for tow buckling to occur. Several numerical works have been completed in this direction. Gatouillat et al. [START_REF] Tephany | Development of an experimental bench to reproduce the tow buckling defect appearing during the complex shape forming of structural flax based woven composite reinforcements[END_REF] proposed a fully meso-scale approach using shell elements in order to predict the tow sliding defects and in [START_REF] Iwata | Coupled meso-macro simulation of woven fabric local deformation during draping[END_REF], Iwata et al. investigated macro-scale and meso-scale numerical simulations to predict the buckling of the tows. They finally proposed the use of a hybrid meso/macro-scale approach. The meso/macro-scale simulation takes advantage of macro-scale simulation to determine shear and tensions which in turn are used at meso-scale as boundary conditions in order to predict buckling and tow displacements; mesoscale simulation is only performed in areas where defect are expected to take place. Local curvature is classically evaluated in numerical The main advantage of using such a framework comes from the improved continuity of the kinematic approximation that can be built in order to be C 2 everywhere in the domain (even at the frontier between two elements) in order to make the curvature continuous. The mechanical behavior of the fabric is taken into account considering the tensile and in-plane shear contributions of the network separately by computing respectively the corresponding matrix, and in the same manner as developed in [START_REF] De Luycker | Simulation of 3D interlock composite preforming[END_REF][START_REF] Hamila | Simulations of textile composite reinforcement draping using a new semi-discrete three node finite element[END_REF]. A full description of the constitutive model is beyond the scope of this paper.

Tangent vectors (defied Eq. ( 22)) are aligned with the material directions of the elements which are chosen to be similar to the fibrous network.

∑

= ∂ ∂ = ∂ ∂ = t M q ξ R ξ (i 1, 2) i i A i A A,p (22)
Normal vectors are defined Eq. ( 23).

∑ = ∂ ∂ = ∂ ∂ = n M q ξ R ξ (i 1, 2) i i A i A 2 2 2 A,p 2 (23)
Using and , it is possible to evaluate the curvature of each tow according to Eq. ( 24):

= ∧ t n t C || || || || i i i i 3 (24)
C i is the total curvature along the i th direction (i = 1,2), computed from the tangent vector and the normal vector which are both obtained from first and second derivatives of the position vector with respect to parameter coordinate ξ i .

In the tow buckling problem, the in-plane bending, seems from previous investigation, to have the most influence for flat tows; the bending along warp and weft directions are computed following tangent ant normal vectors according to first (warp) and second (weft) directions. We define a unit vector normal to ( ) plane:

= ∧ ∧ n t t t t || || 12 1 2 1 2 (25)
The planar component of those normal vectors (associated to the in-plane curvature) is evaluated:

= - n n n n n ( • ) i pl i i 12 12 (26)
Finally, the planar curvature C i pl along the i th direction is:

= ∧ t t C || n || || || i pl i i pl i 3 (27)
Fig. 20(a) shows that zones of low curvatures can be predicted on the forming simulation of a prismatic shape. The highest in-plane curvatures can be observed on the center and the edges of the triangular face of the prism. The curvatures detected in the red zones of Fig. 20(a) are higher than the curvature level from which tow buckles are expected to appear on the three flax and the carbon fabric tested in this work. These zones do correspond to the buckles locations observed in Fig. 20(b) for a very different interlock fabric [START_REF] Ouagne | Analysis of defects during the preforming of a woven flax reinforcement[END_REF].

The zones where tow buckles are expected to appear mainly depend on the shape that induces local curvatures of the tows. However, the magnitude of the tow buckle elevation and the curvature at which the tow buckle may arrive depends on the reinforcement characteristics such as the tow geometry, the tow rigidity and the level of tension of the tows within the shape. The procedure developed in this work therefore permits to indicate according to the fabric characteristics and tension levels in the tows, the zones where tow buckles may take place. It is therefore possible to predict the appearance of tow buckles for a given reinforcement when the constituting tows are experimentally characterized.

Summary and conclusions

In this research work, we focus on the local buckling (at mesoscopic scale) of flat tows. Subjected to in-plane bending, a longitudinal compression is responsible in some configurations for a local buckling. This defect, leading to poor parts quality, was investigated experimentally and analytically.

Experimentally, a dedicated device capable of imposing tensions and an in-plane bending, combined with stereo digital image correlation was used to evaluate the impact of loading, the impact of the tow behavior and the impact of fabric geometry on the onset and kinematic of the buckle growth.

-We found first that an increased stiffness (in tension and in bending) tends to speedup the buckle phenomenon and to increase the high and rotation of the buckles. -Also, fabrics with high unsupported length are more sensitive to the buckle defect (2 × 2 twill exhibits more buckles than plain weave for comparable tow width and yarn densities). -Finally, the experimental study conclusion was that the main driving parameter for the buckling phenomenon was the curvature radius and that tensions impact the buckles through the curvature radius.

Those tendencies were confirmed analytically using a Rayleigh-Ritz approach considering the unsupported portion of a tow in a fabric as an orthotropic plate subjected to bending.

-This approach successfully provide the critical curvature radius for a given fabric constituted of tows of a given nature once the mechanical characteristics of the tows were measured together with the geometrical properties of the fabric.

Results from either the experimental or the analytical approaches can be of valuable interest in a macroscopic simulation of the forming process of a dry fabric. Critical curvature radius can be compared to a simulated one in order to predict the likelihood and location of local buckling defects in a complex part.

-We chose to illustrate this result using an Isogeometrical framework among others for its capability to capture naturally the curvature. -No additional development is needed and only a post processing is required to gather information regarding the buckle onset in a forming process.

This approach is, of course, very simplified since no tow sliding can be taken into account because of the Lagrange description used for the kinematic. Nevertheless, in classical fabric forming, the sliding defect corresponds to quite extreme scenarios and obtaining a tow buckling criterion in a simulation represents a real gain at low cost. Of course, when tensions in the tows reach certain levels and that tow sliding occurs, meso-scale [START_REF] Gatouillat | Meso modelling for composite preform shaping -simulation of the loss of cohesion of the woven fibre network[END_REF] or smarter multi-scale approaches [START_REF] Iwata | Coupled meso-macro simulation of woven fabric local deformation during draping[END_REF] should be favored.

. simulations using neighboring elements [START_REF] Boisse | Simulation of wrinkling during textile composite reinforcement forming. Influence of tensile, in-plane shear and bending stiffnesses[END_REF][START_REF] Mathieu | Enhanced modeling of 3D composite preform deformations taking into account local fiber bending stiffness[END_REF] in order to take into account the bending behavior during a fabric forming. This behavior is proven to be highly influent on the wrinkle's shapes and sizes.

Meso-scale and meso/macro-scale, while being accurate in the prediction of the defect, still require intensive computation resources. Thus, in this work, the use of a macro-scale simulation within an isogeometrical framework [START_REF] Hughes | Isogeometric analysis: CAD, finite elements, NURBS, exact geometry and mesh refinement[END_REF] is proposed.

Considering a set of rational degree p NURBS shape functions R A,p . Let us denote =8(ξ1,ξ2,ξ3), the coordinates in the 3 dimensional parametric space of a considered point; its position vector at a given time t can be obtained considering the control points positions and the value of the NURBS shape functions at as shown in Eq. [START_REF] Kubiak | Static and dynamic buckling of thin-walled plate structures[END_REF]. M (,ξt)= ∑ A R A,p ( )ξq A (t)

Fig. 1 .

 1 Fig. 1. Principle of the tow-buckling device: (a) kinematic diagram of the initial state of the device, (b) kinematic diagram of the bent state of the device, (c) picture of the device in the initial state, (d) picture of the buckles on the surface of the reinforcement. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)

  1(b) and 3(a), assuming that an originally straight tow rotates rigidly. A local way by extracting the curvature radius, using ImageJ® software, from the fabric images of Fig. 3(b).

Fig. 3 .Fig. 4 .

 34 Fig. 3. (a) Bending angle measurement, (b) Tow's curvature radius measurement and representation of the unsupported length L and width b of the tow. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)

Fig. 5 .

 5 Fig. 5. Tow buckling illustration. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)

Fig. 6 .Fig. 7 .

 67 Fig. 6. Evolution of the maximal height and the rotation of the buckling tow as a function of the (a) unsupported length, (b) unsupported width.

Fig. 8 .

 8 Fig. 8. (a) Evolution of the tow maximal height during the buckling of different materials. (b) Evolution of the tow rotation during the buckling of different materials.

Fig. 9 .

 9 Fig. 9. Evolution of the tow maximal height, rotation and curvature radius during the buckling of (a) sized-up flax tows within the fabric as a function of load 1. (b) Carbon fiber tows within the fabric as a function of load 1.

Fig. 10 .

 10 Fig. 10. Evolution of tow height and tow rotation as a function of the in-plane bending angle for a sized-up flax tow under load 1 of 100 N and load 2 of 20 N.

)Fig. 11 .

 11 Fig. 11. Final sized up flax tow profiles under an in-plane bending angle of ~40°.

Fig. 12 .Fig. 13 .

 1213 Fig. 12. Tow Sliding during the tow in-plane bending.

  D11 and D22 are respectively the bending stiffness in the longitudinal and transversal directions, representing the required torque (M x in longitudinal and M y in transversal direction) necessary to generate a unit curvature (K x and K y ). D11 and D22 can therefore be considered as the flexural rigidity (B x and B y ) in a direction normalized by the perpendicular dimension (respectively width: b and length: L).

Fig. 14 .Fig. 15 .

 1415 Fig. 14. Tow representation as an orthotropic plate with compressive load P representation for different values of the non-uniformity coefficient α.

Fig. 16 .

 16 Fig. 16. (a) Peirce test on flax tow (b) Schematic representation of Peirce's cantilever test.

Fig. 17 .

 17 Fig. 17. (a) Torsion test of ARES rheometer®. (b) Torque as a function of unit angle obtained by a torsion test.

Fig. 18 .Fig. 19 .

 1819 Fig. 18. Digital image correlation of the in-plane deformations of a tow bent in its plane before appearance of tow buckling. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)

Fig. 20 .

 20 Fig. 20. (a) In-plane curvature radii along x (left) and y (right) fiber networks directions (logscale) during a prismatic fabric forming. (b) Location of buckles during the deep drawing of a carbon interlock fabric on a prismatic shape [13] the initial direction of the fabric is aligned along the edges. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)

Table 2

 2 Tows dimensional and mechanical properties with analytical and experimental critical buckling radii.

	Tow origin	Twill	Twill	Twill	Plain A	Plain C
		2 × 2	2 × 2	2 × 2	desized	desized
		sized up	desized	carbon	up flax	up flax
		flax	up flax	fiber
	Linear density (g/	0.614	0.494	0.813	0.488	0.478
	m)
	E1 (N/mm 2 )	19,800	7900	132,700	7820	7645
	L (mm)	6.11	5.66	6.42	6.46	2.83
	b (mm)	2.48	2.40	2.6	2.10	2.65
	h (mm)	0.4	0.4	0.38	0.4	0.4
	α	2	2	2	2	2
	m	1	1	1	1	1
	D11 (N•mm)	1.18	0.72	7.55	0.72	0.63
	D66 (N•mm)	0.44	0.17	-	0.17	0.17
	Rcr analytical (mm) 2082.9	1482.2	-	1275.7	717.7
	Rcr experimental	2255	1627	-	1390	803
	(mm)
	Δ (%)	7.63	8.90	-	8.22	10.62


C Eline Crauste

Christian P Erigaud

Suzanne Peyrottes
email: suzanne.peyrottes@univ-montp2.fr

Insights into the Soluble PEG-Supported Synthesis of Cytosine-Containing Nucleoside 5 0 -Mono-, Di-, and Triphosphates

Insights into the Soluble PEG-Supported Synthesis of Cytosine-Containing Nucleoside 5'-Mono-, Di-, and Triphosphates

Céline Crauste, Christian Périgaud, Suzanne

Introduction

Nucleotides, especially nucleoside 5 0 -phosphorylated forms, are involved in many fundamental biological processes and pathways, 1 and constitute an important area of research. Thus, several procedures have been reported for the conversion of a nucleoside to its corresponding 5 0 -mono-, di-, or triphosphorylated forms, using classical solution synthesis. 2 The phosphorylation success depends considerably on the nature of the substrate and to date no protocol for making nucleoside 5 0 -triphosphates is universally satisfactory. Moreover, solution methods usually yield to complex crude mixtures which require tedious purification steps involving first ion-exchange chromatography on DEAE-Sephadex in aqueous buffers. In addition to the desired nucleotides, the crude mixture also contains large amounts of inorganic salts, such as inorganic phosphates, that may be eliminated by an additional purification step by using preparative C18 reverse-phase HPLC. Usually, this fastidious and long purification procedure, associated with the relative low chemical stability of these derivatives, induces a decrease in yields.

Since Merrifield et al. 3 introduced the use of polymersupported synthesis of oligopeptides, insoluble supports have become an important tool for organic synthesis, 4 especially for the synthesis of biopolymers, 5,6 such as oligonucleotides, peptides, and carbohydrates. Few works have been reported on the supported synthesis of nucleoside triphosphates, [7][8][9][10][11][12][13][14][15][16][17][18] with the aim of reaching completion of the reaction (use of excess reagent) and to greatly simplify purification processes. These works were based on the use of the phosphoramidite chemistry and among these, only Parang et al. [7][8][9][10][11][12][13] have developed a regioselective method for nucleoside 5 0 -phosphates and 5 0 -thio-phosphates synthesis. Thus, the use of a new phosphorylation reagent immobilized on solid support allowed the desired derivatives to be isolated in good yields (∼80%) and acceptable HPLC purity (70-90%).

Despite the benefits offered by the solid phase approach, the heterogeneous reaction conditions (due to the nature of the insoluble polymers) often induce other problems such as nonlinear reaction kinetics, unequal distribution and/or access to the reaction sites, solvation problems, inefficient coupling rates, and difficulties for intermediates analysis. The search for alternative methodologies, aiming to restore homogeneous reaction conditions, has led to the development of soluble polymer supports. These last have received considerable attention because such a "liquid phase" synthesis retains many of the conventional solution chemistry advantages (high reactivity, lack of diffusion, and easy analysis) while still allowing a smooth isolation and purification of the product. 19,20 In addition, soluble polymers are usually much less expensive than insoluble ones. As an example, soluble polyethylene glycol (PEG), polyvinyl alcohol, and other polymers have been successfully used for the synthesis of bio-oligomers (peptides, 21,22 oligonucleotides, 23,24 polysaccharides) 25 and small molecules. 19 More recently fluorous phase [26][27][28] and ionic liquids [29][30][31][32][33][34][35] have also been reported as useful soluble supports for organic synthesis and olignucleotide preparation. 36,37 Among these soluble polymer supports, PEG has been wildly used for the liquid-phase oligonucleotide synthesis. 38 Borona et al. [39][40][41][42][START_REF] Bonora | Nucleosides[END_REF][44][45][46][47] have developed a high-efficiency liquid phase (HELP) method for large-scale synthesis of oligonucleotides based on a PEG strategy; to date these works have not been adapted for nucleotide (nucleoside mono-, di-, and triphosphates) preparation. Thus, to develop a new nucleotide synthetic approach and to easily transfer the traditional solution-phase conditions to a supported methodology, we propose herein a new technology based on the use of a soluble PEG support.

Results and Discussion

Supported Nucleotide Synthesis: Polyethylene Glycol (PEG). The wide application of PEG in organic synthesis is closely related to its broad solubility profile. It is soluble in N,N-dimethylformamide (DMF), dichloromethane, toluene, acetonitrile, water, and methanol, but insoluble in diethyl ether, tert-butyl methyl ether, isopropyl alcohol, and cold methanol. Therefore, excess reagents and some byproduct could readily be eliminated at the end of each reaction step by using a simple precipitation from diethyl ether of the polymer-substrate followed up by filtration, and eventually additional purification by flash chromatography may be performed. A commercially available PEG of 4000 Da average molecular weight was selected as the starting material. Analyses of each supported intermediates were performed by nuclear magnetic resonance spectra ( 1 H, 13 C, and 31 P NMR).

Functionalization of PEG and Coupling

Step. The synthetic strategy is based on PEG anchored to the nucleoside (nucleobase) through a basolabile linker. This procedure was selected to develop a synthetic method suitable for the preparation of nucleotide modified on the glycosidic moiety; indeed such a modification is often encountered within nucleoside analogues with antiviral and antitumoral properties. Thus, we focused on the synthesis of nucleotides containing cytosine as nucleobase namely cytidine (C), 2 0 -deoxycytidine (dC), and cytarabine (araC, 1-β-D-arabinofuranosylcytosine), a nucleoside analogue commonly used as an antitumoral agent in the treatment of acute myelogenous leukemia. 48,49 Initial functionalization of PEG support with the basolabile linker was performed through the coupling of PEG-(OH) 2 with succinic anhydride in the presence of a catalytic amount of 4-dimethylaminopyridine (DMAP) in pyridine. 45 A quantitative rate of fonctionalization was estimated by 13 C NMR analysis by observing the shifts of CH 2 R and CH 2 β PEG signals (Figure 1).

Then, the anchoring of C, dC, and araC was performed as follows. Each nucleoside was linked to the PEG-(Osuccinate) 2 1 through the exocyclic amino group of the JOCArticle nucleobase in the presence of 1-hydroxybenzotriazole (HOBT) and N,N 0 -dicyclohexylcarbodiimide (DCC) as coupling agents in a mixture of anhydrous dichloromethane and DMF (Scheme 1). 50 This protocol allowed anchoring regioselectively the selected nucleoside onto the polymer without prior protection of the sugar hydroxyl functions. The PEGnucleosides 2a, 2b, and 2c were unequivocally characterized by 1 H NMR spectroscopy, which also confirms the complete removal of reagents (first ascertained by TLC). Coupling rates were estimated by comparing 1 H NMR integration of both signals of the H 6 from the nucleobase and of the methylene groups of the succinyl linker (Figure 2). Conversions were quantitative and yields were between 83% and 87%. With one nucleoside at each end of the polymeric chain, the capacity of the synthesized PEG-nucleoside was evaluated to be 0.5 mmol 3 g -1 .

Monophosphorylation Step. In a preliminary set of experiments, the monophosphorylation reaction was optimized with araC as nucleosidic model. To prepare nucleoside 5 0 -monophosphates the regioselective method of Yoshikawa et al., 52,53 previously described for solution synthesis, was adapted. As conventional monophosphorylation of unprotected nucleosides requires the use of trialkylphosphate as solvent and phosphorus oxychloride as phosphorylating agent, an initial attempt was carried out in these conditions. Unfortunately, PEG-araC 2a was soluble at neither 0 °C nor room temperature in triethylphosphate. The solubility of the PEGsupported nucleoside was only achieved when reaching the temperature of 35-40 °C. As acetonitrile could solubilize the PEG moiety at 0 °C, it was also considered an alternative solvent. Thus, several experimental conditions (solvent, temperature, reaction times, and phosphorus oxychloride equivalents) were screened and selected data are presented in Table 1. Monophosphorylation conversion ratios were estimated with 1 H NMR analysis by comparing integration of the signals of the H 6 or H 5 nucleobase protons (Figure 3). Indeed, the signals of the H 6 proton from the nucleoside 5 0 -monophosphate differed from that of the nucleoside by 60 Hz (0.2 ppm, downfield).

These conversion ratios were also confirmed with HPLC analysis of a sample aliquot after cleavage of the succinyl linker in between the support and the nucleot(s)ide. In each case, regioselectivity of the phosphorylation (at the 5 0 -position) was evidenced by the 1 H NMR and 31 P NMR spectra. Longer reaction times did not improve the phosphorylation rate significantly (entries 5-7) whereas increasing the phosphorus oxychloride amount led to the reaction completion (entries 2-4 and 5, 8, and 11). Thus, the use of dry triethylphosphate appeared a better alternative than acetonitrile (entries 1 and 10). As both solvents were able to dissolve the substrate (PEG-araC) and the reactant (POCl 3 ), the effect of the triethylphosphate was attributed to the formation of a complex 54 between the nucleobase and the 5 0 -hydroxyl function of the nucleoside, and not to solubility or temperature effects (as previously reported for solution phase synthesis). 52,53 Finally, PEG-araCMP 3a was obtained by treating PEG-araC 2a with 15 equiv of POCl 3 in PO(OEt) 3 at 35-40 °C during 1 h. After hydrolysis of the reaction mixture with TEAB (1M, pH 7.5), remaining salts were eliminated by a simple and easy extraction process (dichloromethane-water) and PEG-supported nucleoside 5 0 -monophosphate 3a was isolated in 92% yield, after precipitation in cold diethyl ether.

Recovery of araCMP (HPLC purity: 97%) was performed after cleavage of the succinyl linker, using concentrated ammonia. A filtration of the crude through a reverse phase RP 18 column was performed to eliminate PEG residue and succinimide (the byproduct generated from the cleavage of succinate linker) was removed by dialysis.

A similar procedure was applied to PEG-C 2c and the use of 30 equiv of POCl 3 was required to observe the complete disappearance of the supported nucleoside. During the PEG-dCMP 3b synthesis, the 3 0 -hydroxyl function also underwent phosphorylation at high temperature (30-40 °C) and a mixture of both nucleoside 5 0 -and 3 0 -monophosphates was obtained. To avoid this lack of regioselectivity the reaction was performed in acetonitrile at 0 °C. At lower temperature, only the desired nucleoside 5 0 -monophosphate 3b was observed and the use of 30 equiv of POCl 3 was necessary to complete the reaction. Finally, CMP and dCMP were obtained after ammonia treatment; corresponding yields and purities are presented in Scheme 1.

JOCArticle

Di-and Triphosphorylation Steps. The triphosphorylation procedure was first considered following the Ludwig et al. 55 one-pot method, using unprotected nucleoside as starting material. Consequently, the phosphorodichloridate intermediate was generated with use of the monophosphorylation conditions previously optimized (PO(OEt) 3 , 35-40 °C, 15 equiv of POCl 3 , 1 h) and the reaction mixture was treated by an excess of tri-n-butylammonium pyrophosphate (1 M solution in DMF). Unfortunately, hydrolysis and purification afforded a mixture of the mono-and the triphosphate-supported nucleosides with a prevalent proportion of PEG-araCMP 3a. As the completion of the reaction was a key factor for success, another method described by Hoard et al., 56 which could easily be monitored by 31 P NMR analysis, was investigated. This procedure required the nucleoside 5 0 -monophosphate activation by 1,1-carbonyldiimidazole (CDI) and then condensation of the 5 0 -phosphoroimidazolidate intermediate with inorganic phosphate or pyrophosphate, which led to the desired nucleoside 5 0 -di-or triphosphate derivatives (Scheme 2).

Formation of the imidazolidate from PEG-araCMP 3a in the presence of an excess of CDI (6 equiv) was complete after 3 h. Indeed, the 31 P NMR signal corresponding to the supported nucleoside monophosphate (signal at 0 ppm, Figure 4A) disappeared whereas the concomitant formation of the imidazolidate (signal at -10 ppm, Figure 4B) was observed. Unreacted CDI was decomposed with methanol and then a large excess of organic phosphate was added. As shown in Figure 4C, 31 P NMR signals observed after 24 h confirmed that activated imidazolidate species (δ -10 ppm) were efficiently converted to the diphosphate derivative (two peaks corresponding to P β at -9.2 ppm and to P R at -12 ppm). As the precipitation and extraction processes were not sufficient to eliminate the excess phosphate salts, purification of the PEG-supported nucleoside diphosphate 5a was performed by reverse phase C18 column chromatography. Elution with water removed all remaining salts and supported substrate was eluted with a mixture of water and acetonitrile (50/50, v/v). The crude supported product was lyophilized to afford PEG-araCDP 5a in 66% yield. A similar procedure was applied for PEG-araCTP 6a synthesis (69%), using tri-nbutylammonium pyrophosphate salts instead of tri-n-butylammonium phosphate salts. Then, treatment of 5a and 6a with concentrated aqueous ammonia yielded araCDP 7a and araCTP 8a, as ammonium salts, respectively, with 87% and 83% HPLC purity after RP18 purification and dialysis. Starting from PEG-dCMP 4b, the same procedures were carried out and afforded either the dCDP 7b or the dCTP 8b (Scheme 2).

When applied to PEG-CMP 3c, these protocols led to the expected derivatives PEG-CDP 5c and PEG-CTP 6c as well as two side products (Scheme 3). The major contaminant was identified as the corresponding 2 0 ,3 0 -protected nucleotide (9 or 10). Formation of these derivatives is due to the particular reactivity of the cis-diol functionality of ribosecontaining nucleosides in the presence of CDI excess. However, formation of these derivatives was not puzzling as concomitant removal of the 2 0 ,3 0 -carbonate group was expected during the final cleavage step with basic conditions.

Structures of minor contaminants 11 and 12 (dioxaphospholane derivatives involving a cyclic monophosphate between the 2 0 and the 3 0 positions) were proposed on the basis of the presence of an extra phosphate signal at δ 19 ppm, already described in the literature for 2 0 ,3 0 -cyclic phosphate nucleosides. 57,58 Furthermore, the HPLC-MS analysis of the crude mixture after cleavage showed peaks corresponding to compounds 13 and 14. In the meantime, El Tayeb et al. 58 reported that phosphorylation of the 2 0 -and 3 0 -hydroxy groups may occurred during the triphosphorylation, 55 and this side reaction is associated with the presence of bulky substituents on the pyrimidine nucleobase. In our case, steric hindrance of PEG support attached onto the nucleobase may be involved in the 2 0 ,3 0 phosphorylation process. Indeed, such a side reaction was not observed when the reaction was performed in solution from unprotected cytidine 5 0 -monophosphate (cf. CTP solution synthesis).

Thus, after performing the cleavage step CDP 7c or CTP 8c was obtained, albeit contaminated by a small amount of 2 0 ,3 0 -cyclic phosphate 13 or 14 (7% and 10%, respectively) and unfortunately final purification on reverse phase did not allow the separation of this byproduct.

Comparison between Solution and Supported Nucleotide Synthesis. To compare our novel soluble PEG-supported methodology to the corresponding synthesis in solution (in terms of number of step, yields, and HPLC purities, Table 2) we also prepared araCMP, araCDP, and araCTP using traditional methods.

AraCMP was obtained by solution phase synthesis, using phosphorus oxychloride in PO(OEt) 3 as previously reported by Yoshikawa. 52,53 After cumbersome purification steps (ion exchange chromatography DEAE-Sephadex, reverse phase C18 chromatography, and conversion into sodium salt with Dowex resin) araCMP was recovered in modest to good yields depending of the reaction scale (70% for 200 mg of

JOCArticle

starting material and 40% for 1 g). This result is mainly due to the numerous purification steps needed to eliminate the remaining salt on a larger scale. Supported monophosphorylation was also carried out on two different scales (200 mg and 1 g of supported nucleoside), meanwhile no variation on the overall yields was observed (60% including coupling of the nucleoside to the support, monophosphorylation, and cleavage steps). Solution synthesis of araCDP and araCTP was also performed following the Hoard et al. 56 method starting from araCMP. For the araCDP synthesis yields on both support and in solution were similar (∼50%). Remarkably, using our supported methodology the yield of the nucleoside 5 0 -triphosphate synthesis was 5-fold the one obtained by the solution approach.

One can see that we also succeeded in reaching similar HLPC purities with 5 0 -mono-, di-, and triphosphate nucleoside-supported synthesis (Table 2), without performing ion-exchange DEAE-Sephadex purification but only using precipitation/ extraction or precipitation/RP18 filtration processes.

Conclusion

We have set up a new synthetic approach for the supported synthesis of nucleoside phosphoesters using PEG as a soluble support. Compared to classical solution procedures, our methodology is efficient, scalable, and easy to handle. Thus, cytosine-containing nucleoside 5 0 -mono, di-, and triphosphate derivatives were prepared in similar higher yields with acceptable purity. Furthermore, the isolation procedures have been simplified and led to a reduction in the overall process time.

a

  SCHEME 1 a

FIGURE 1 .

 1 FIGURE 1. Determination of PEG functionalization with succinyl linker by 13 C NMR (100 MHz) analysis: (left) starting PEG-(OH) 2 13 C NMR spectrum (CDCl 3 ) with δ 61.5 ppm for CH 2 R and δ 72.8 ppm for CH 2 β; (right) PEG-(O-succinate) 2 13 C NMR spectrum (CDCl 3 ) with δ 64.2 ppm for CH 2 R and δ 68.9 ppm for CH 2 β.

FIGURE 2 .

 2 FIGURE 2. 1 H NMR (300 MHz) spectrum of PEG-C (D 2 O).

FIGURE 4 .

 4 FIGURE 4. 31 P NMR (DMSO-d 6 ) spectra of the PEG-araCMP diphosphorylation procedure: (A) starting material PEG-araCMP; (B) reaction mixture, 3 h after CDI addition; and (C) reaction mixture 24 h after tri-n-butylammonium phosphate addition. SCHEME 2 a

a

  SCHEME 3 a

TABLE 1 .

 1 Optimization of Supported araC Monophosphorylation Step Determined by HPLC 51 after cleavage of the linker in between the support and the nucleot(s)ide. Bull. Chem. Soc. Jpn. 1967, 40, 2849-2853. (53) Yoshikawa, M.; Kato, T.; Takenish, T Tetrahedron Lett. 1967, 5065-5068.

	entry	equiv ofPOCl 3	time (h)	solvent/temp (°C)	% HPLC a araC/araCMP
	1	1 0	6	C H 3 CN/0	53	47
	2	2 0	3	C H 3 CN/0	31	69
	3	3 0	3	C H 3 CN/0	24	76
	4	5 0	3	CH 3 CN/0	18	82
	5	5	1	PO(OEt) 3 /35-40	57	43
	6	5	3	PO(OEt) 3 /35-40	69	31
	7	5	6	PO(OEt) 3 /35-40	69	31
	8	10	1	PO(OEt) 3 /35-40	15	85
	9	10	3	PO(OEt) 3 /35-40	14	86
	10	10	6	PO(OEt) 3 /35-40	11	89
	11	15	1	PO(OEt) 3 /35-40	3	9 7
	12	15	3	PO(OEt) 3 /35-40	9	91
	13	15	6	PO(OEt) 3 /35-40	7	93
	a

FIGURE 3. Determination of monophosphorylation rate by 1 H NMR analysis: 1 H NMR (300 MHz, in D 2 O) spectrum obtained after monophosphorylation of PEG-araC in experimental conditions corresponding to entry 1, Table 1

.

(52) Yoshikawa, M.; Kato, T. (54) Ikemoto, T.; Haze, A.; Hatano, H.; Kitamoto, Y.; Ishida, M.; Nara, K. Chem. Pharm. Bull. 1995, 43, 210-215.

TABLE 2 .

 2 Comparison of Solution and Supported Synthesis of Phosphoester Derivatives of araC Overall yield for supported synthesis included coupling, phosphorylation, and cleavage and isolation steps. (57) Davies, D. B.; Sadikot, H. J. Chem. Soc., Perkin Trans. 2 1983, 1251-1258. (58) El-Tayeb, A.; Qi, A. D.; Muller, C. E. J. Med. Chem. 2006, 49, 7076-7087.

	solution

Acknowledgment. C.C. is grateful to the French Ministry of Education and Research (MENRT) for a Ph.D. grant and we thank the "Centre National de Recherches Scientifiques" (CNRS) for funding.

Experimental Section

General Procedure for PEG-Supported Nucleoside Synthesis. To a stirred solution of PEG succinate 1 (0.95 mmol, 1 equiv) in dichloromethane (70 mL) was added a solution of the required nucleoside (2 equiv) in DMF (30 mL). Then, DCC (3.79 mmol, 4 equiv) and HOBT (1.89 mmol, 2 equiv) were added and the resulting mixture was stirred at 60 °C for 7 h. The solvents were evaporated under reduced pressure; the residue was dissolved in a minimum of dichloromethane and left overnight at 4 °C. The dicyclohexyl urea was filtered off, and the filtrate was concentrated under reduced pressure and dissolved in DMF (40 mL). The resulting solution was slowly added to an excess volume of cold diethyl ether (400 mL). The precipitate was filtered and washed with diethyl ether. The final product was recrystallized from absolute ethanol (40 mL) and dried under vacuum over KOH pellets.

Poly(ethylene glycol) 4000 bis[4-N-(1-(β-D-arabinofuranosyl)cytosyl)succinate] 2a: Coupling between PEG succinate 1 (4 g, 0.95 mmol) and araC (0.46 g, 1.89 mmol) was performed with the general procedure and afforded compound 2a as a white solid (3.65 g, 83%). R f (CH 2 Cl 2 /MeOH, 7/3, v/v) 0.7; δ H (200 MHz; D 2 O) 8.16 (d, J 6-5 = 7.5 Hz, 1H, H 6 ), 7.26 (d, J 5-6 = 7.5 Hz, 1H, H 5 ), 6.10 (d, J 1 0 -2 0 = 4.6 Hz, 1H, H 1 0 ), 4.34 (pt, J = 4.6 Hz, 1H, H 2 0 ), 4.16 (m, 2H, (OCH 2 R) PEG ), 4.02-3.92 (m, 2H, H 3 0 H 4 0 ), 3.73 (m, 2H, H 5 0 a H 5 0 b ), 3.70-3.33 (m, (OCH 2 ) PEG ), 2.76-2.62 (m, 4H, CH 2succ ); δ C (75 MHz; D 2 O) 174.6, 174.5 (2s, CdO succ ), 162.5 (s, C 4 ), 156.7 (s, C 2 ), 146.7 (s, C 6 ), 97.3 (s, C 5 ), 87.1 (s, C 1 0 ), 84.0 (s, C 4 0 ), 75.4 (s, C 2 0 ), 75.2 (s, C 3 0 ), 69.6 (s, (OCH 2 ) PEG ), 68.4 (s, (OCH 2 β) PEG ), 64.1 (s, (OCH 2 R) PEG ), 60.7 (s, C 5 0 ), 31.5, 28.5 (2s, CH 2succ ).

General Procedure for the Monophosphorylation of Polymer-Supported Nucleoside. A solution of PEG-supported nucleoside (1 equiv) in triethylphosphate (10 mL/mmol) was warmed to 40 °C. Phosphorus oxychloride (15 equiv) was added and the mixture was stirred for 1 h at 40 °C. Excess of reagent was hydrolyzed by addition of an aqueous triethylammonium bicarbonate buffer solution (1M, pH 7) and the resulting mixture was concentrated under reduced pressure. The residue was dissolved in dichloromethane and the organic layer was washed with water. The aqueous layer was extracted several times by dichloromethane. The organic layers were combined and evaporated under reduced pressure. The supported nucleoside 5 0 -monophosphate was precipitated from a dichloromethane solution by addition of an excess volume of cold diethyl ether. The precipitate was filtered, washed with diethyl ether, and dried under vacuum over KOH.

Poly(ethylene glycol) 4000 bis[4-N-(1-(5 0 -O-monophosphoryl-β-D-arabinofuranosyl)cytosyl)succinate], triethylammonium salt 3a: Monophosphorylation of PEG-araC 2a (3 g, 0.64 mmol) was performed with the general procedure and afforded compound 3a as a white solid (2.99 g, 92%). δ H (300 MHz; D 2 O) 8.34 (d, J 6-5 = 7.5 Hz, 1H, H 6 ), 7.10 (d, J 5-6 = 7.5 Hz, 1H, H 5 ), 6.18 (d, J 1 0 -2 0 = 5.0 Hz, 1H, H 1 0 ), 4.42 (m, 1H, H 2 0 ), 4.19 (m, 2H, (OCH 2 R) PEG ), 4.10-4.15 (m, 4H, H 3 0 H 4 0 H 5 0 a H 5 0 b ), 3.30-3.90 (m, (OCH 2 ) PEG ), 3.10 (q, J = 7.3 Hz, 6H, (CH 3 CH 2 ) 3 NH), 2.62-2.85 (m, 4H, CH 2succ ), 1.18 (t, J = 7.3 Hz, 9H, (CH 3 -CH 2 ) 3 NH); δ C (100 MHz; D 2 O) 174.7, 174.4 (2s, CdO succ ), 159.0 (s, C 4 ), 148.1 (s, C 2 ), 145.2 (s, C 6 ), 96.9 (s, C 5 ), 87.0 (s, C 1 0 ), 82.3 (d, J C4 0 -P = 8.0 Hz, C 4 0 ), 75.8 (s, C 2 0 ), 74.1 (s, C 3 0 ), 69.6 (s, (OCH 2 ) PEG ), 68.4 (s, (OCH 2 β) PEG ), 64.1 (s, OCH 2 R) PEG ), 63.6 (s, C 5 0 ), 46.6 (s, (CH 3 CH 2 ) 3 NH), 31.6, 28.3 (2s, CH 2succ ), 8.2 (s, (CH 3 CH 2 ) 3 NH); δ P (121 MHz; D 2 O) 0.26 (s).

General Procedure for the Diphosphorylation of Polymer-Supported Nucleoside. A suspension of PEG-supported nucleoside 5 0 -monophosphate (1 equiv) in tributylamine (10 mmol, 100 equiv) was stirred for 10 min at room temperature. Cold diethyl ether was added and the precipitate was filtered, washed with diethyl ether, and dried overnight under vacuum over KOH pellets. Then, the supported nucleoside 5 0 -monophosphate was dissolved in anhydrous DMF (10 mL/mmol) and CDI (6 equiv) was added. The reaction mixture was stirred for 2 h at room temperature and treated with anhydrous methanol (24 equiv) to hydrolyze the excess of CDI. After 15 min, a solution of tributylammonium phosphate (1 M solution in anhydrous DMF, 15 equiv) was added and the suspension was stirred for 24 h at room temperature. The mixture was treated with an equal volume of methanol and then concentrated under reduced pressure. The residue was precipitated with cold diethyl ether; the precipitate was filtered and washed with diethyl ether. Purification of the crude material was performed on reverse phase column (linear gradient from water to 70% acetonitrile in water).

Poly(ethylene glycol) 4000 bis[4-N-(1-(5 0 -O-diphosphoryl-β-Darabinofuranosyl)cytosyl)succinate], tributylammonium salt 5a: Diphosphorylation of PEG-araCMP 3a (0.50 g, 0.09 mmol) was performed with the general procedure and afforded compound 5a as a white solid (0.36 g, 66%) after freeze-drying. δ H (300 MHz; D 2 O) 8.24 (d, J 6-5 = 7.5 Hz, 1H, H 6 ), 7.27 (d, J 5-6 = 7.5 Hz, 1H, H 5 ), 6.17 (d, J 1 0 -2' = 5.1 Hz, 1H, H 1 0 ), 4.42 (pt, J = 5.1 Hz, 1H, H 2 0 ), 4.28-4.05 (m, 6H, H 3 0 H 4 0 H 5 0 a H 5 0 b (OCH 2 R) PEG ), 3.89-3. 35

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General Procedure for the Triphosphorylation of Polymer-Supported Nucleoside 5 0 -Monophosphate. A suspension of PEG-supported nucleoside 5 0 -monophosphate (1 equiv) in tributylamine (10 mmol, 100 equiv) was stirred for 10 min at room temperature. Then, cold diethyl ether was added and the precipitate was filtered, washed, and dried overnight under vacuum over KOH pellets. The resulting supported nucleoside 5 0 -monophosphate was dissolved in anhydrous DMF (10 mL/mmol) and CDI (6 equiv) was added. The anhydrous reaction mixture was stirred for 2 h at room temperature and treated with anhydrous methanol (24 equiv) to hydrolyze the CDI excess. After 15 min, a solution of tributylammonium pyrophosphate (1 M solution in anhydrous DMF, 15 equiv) was added. The suspension was stirred for 24 h at room temperature. The mixture was treated with an equal volume of methanol and then concentrated under reduced pressure. The residue was precipitated with cold diethyl ether. The precipitate was filtered and washed with diethyl ether. Column chromatography of the crude material was performed on reverse phase (linear gradient from water to 70% acetonitrile in water).

Poly(ethylene glycol) 4000 bis[4-N-(1-(5 0 -O-triphosphoryl-β-Darabinofuranosyl)cytosyl)succinate], tributylammonium salt 6a: Triphosphorylation of PEG-araCMP 3a (0.50 g, 0.09 mmol) was performed with the general procedure and afforded compound 6a as a white solid (0.43 g, 69%) after freeze-drying. δ H (300 MHz; D 2 O) 8.29 (d, J 6-5 = 7.5 Hz, 1H, H 6 ), 7.17 (d, J 5-6 = 7.5 Hz, 1H, H 5 ), 6.17 (d, J 1 0 -2' = 5.1 Hz, 1H, H 1 0 ), 4.41 (pt, J = 5.1 Hz, 1H, H 2 0 ), 4.27-4.05 (m, 6H, H 3 0 H 4 0 H 5 0 a H 5 0 b (OCH 2 R) PEG ), 3.88-3.30 (m, (OCH 2 ) PEG ), 3.03 (t, J=7.4 Hz, 18H, (CH 3 CH 2 CH 2 CH 2 ) 3 NH), 2.87-2.65 (m, 4H, CH 2succ ), 1.57 (m, 18H, (CH 3 CH 2 CH 2 CH 2 ) 3 NH), 1.27 (m, 18H, (CH 3 -CH 2 CH 2 CH 2 ) 3 NH), 0.83 (t, J = 7.1 Hz, 27H, (CH 3 CH 2 CH 2 -CH 2 ) 3 NH); δ C (100 MHz; D 2 O) 174.7, 174.5 (2s, CdO succ ), 161.5 (s, C 4 ), 155.1 (s, C 2 ), 147.6 (s, C 6 ), 97.2 (s, C 5 ), 86.6 (s, C 1 0 ), 82.0 (d, J C4 0 -P = 9.0 Hz, C 4 0 ), 75.2 (s, C 2 0 ), 74.0 (s, C 3 0 ), 69.6 (s, (OCH 2 ) PEG ), 68.4 (s, (OCH 2 β) PEG ), 64.2 (s, (OCH 2 R) PEG ), 64.1 (d, J C5 0 -P = 8.0 Hz, C 5 0 ), 52.6 (s, (CH 3 CH 2 CH 2 CH 2 ) 3 NH), 31.6, 28.4 (2s, CH 2succ ), 25.2 (s, (CH 3 CH 2 CH 2 CH 2 ) 3 NH), 18.3 (s, CH 3 CH 2 CH 2 CH 2 ) 3 NH), 12.8 (s, (CH 3 CH 2 CH 2 -CH 2 ) 3 NH); δ P (121 MHz, D 2 O) -10.94 (d, J γ-β = 19.4 Hz, P γ ), -11.32 (d, J R-β = 19.4 Hz, P R ), -23.28 (t, J β-γ =J β-R = 19.4 Hz, P β ).

General Procedure for the Cleavage of Supported Substrates and Nucleotides Recovery. PEG-supported nucleotide (0.05 mmol, 1 equiv) was dissolved in concentrated ammonia (2 mL). The solution was stirred at room temperature for 1 h and then evaporated under reduced pressure. Purification was performed on reverse phase column (elution with water) and dialysis.

1-(β-D-Arabinofuranosyl)cytosine-5 0 -monophosphate, ammonium salt 4a: The general procedure was applied to PEG-araCMP 3a (0.15 g, 0.02 mmol) and led to a mixture (15 mg) of araCMP and succinamide (ratio: 93/7). The succinamide was eliminated by dialysis in water (2 mL) overnight and araCMP 4a (13 mg, 68%) was recovered as a white solid after freeze-drying. λ max (H 2 O)/272 nm (ε/dm 3 mol -1 cm -1 7100), 247 (4100); δ H (300 MHz; D 2 O) 7.88 (d, J 6-5 = 7.6 Hz, 1H, H 6 ), 6.11 (d, J 1 0 -2 0 = 5.3 Hz, 1H, H 1 0 ), 6.02 (d, J 5-6 = 7.6 Hz, 1H, H 5 ), 4.34 (pt, J = 5.3 Hz 1H, H 2 0 ), 4.11-3.96 (m, 4H, H 3 0 H 4 0 H 5 0 a H 5 0 b ); δ C (100 MHz; D 2 O) 161.2 (s, C 4 ), 151.0 (s, C 2 ), 144.4 (s, C 6 ), 94.7 (s, C 5 ), 85.5 (s, C 1 0 ), 81.5 (d, J C4 0 -P = 9.0 Hz, C 4 0 ), 75.3 (s, C 2 0 ), 73.5 (s, C 3 0 ), 63.1 (d, J C5 0 -P = 4.0 Hz, C 5 0 ); δ P (121 MHz, D 2 O) 0.45 (s); LC/MS (ESI) m/z 322 (M -NH 4 þ ). 1-(β-D-Arabinofuranosyl)cytosine-5 0 -diphosphate, ammonium salt 7a: The general procedure was applied to PEG-araCDP 5a (0.34 g, 0.05 mmol) and led to a mixture (55 mg) of araCDP and succinamide (ratio: 88/12). An aliquot (10 mg) was dissolved in water (2 mL) and submitted to dialysis. AraCDP 7a (8.50 mg, 96%) was obtained as a white solid after freeze-drying. λ max (H 2 O)/274 nm (ε/dm 3 mol -1 cm -1 7900), 246 (3800); δ H (300 MHz; D 2 O) 7.85 (d, J 6-5 = 7.6 Hz, 1H, H 6 ), 6.14 (d, J 1 0 -2 0 = 5.4 Hz, 1H, H 1 0 ), 6.03 (d, J 5-6 =7.6 Hz, 1H, H 5 ), 4.35 (pt, J = 5.4 Hz, 1H, H 2 0 ), 4.20-4.10 (m, 3H, H 3 0 H 5 0 a H 5 0 b ), 4.00 (m, 1H, H 4 0 ); δ C (100 MHz; D 2 O) 166.1 (s, C 4 ), 157.4 (s, C 2 ), 142.7 (s, C 6 ), 95.7 (s, C 5 ), 85.1 (s, C 1 0 ), 81.1 (d, J C4 0 -P = 9.0 Hz, C 4 0 ), 75.3 (s, C 2 0 ), 73.6 (s, C 3 0 ), 63.6 (d, J C5 0 -P = 5.0 Hz, C 5 0 ); δ P (121 MHz, D 2 O) -10.42 (d, J β-R = 20.6 Hz, P β ), -11.10

). 1-(β-D-Arabinofuranosyl)cytosine-5 0 -triphosphate, ammonium salt 8a: The general procedure was applied to PEG-araCTP 6a (0.43 g, 0.07 mmol) and led to a mixture (68 mg) of araCTP and succinamide (ratio: 95/5). An aliquot (8.5 mg) was dissolved in water (2 mL) and submitted to dialysis. AraCTP 8a (8 mg, 92%) was obtained as a white solid after freeze-drying. λ max (H 2 O)/273 nm (ε/dm 3 mol -1 cm -1 8700), 246 (4500); δ H (300 MHz; D 2 O) 7.85 (d, J 6-5 = 7.6 Hz, 1H, H 6 ), 6.14 (d, J 1 0 -2 0 = 5.4 Hz, 1H, H 1 0 ), 6.04 (d, J 5-6 = 7.6 Hz, 1H, H 5 ), 4.35 (pt, J = 5.4 Hz, 1H, H 2 0 ), 4.20-4.12 (m, 3H, H Supporting Information Available: General information, experimental details, and characterization data for reported compounds. This material is available free of charge via the Internet at http://pubs.acs.org.